NIST

PUBLICATIONS

NIST Special Publication 1000-5

June 2004

Progress Report on the

Federal Building and Fire Safety

Investigation of the

World Trade Center Disaster

National Institute of Standards and Technology • Technology Administration • U.S. Department of Commerce

Volume 5

Contains Appendices J, K, L, and M

V.5
4
NIST Special Publication 1000-5

June 2004

Progress Report on the

Federal Building and Fire Safety

Investigation of the

World Trade Center Disaster

Volume 5

Contains Appendices J, K, L, and M

June 2004

U.S. Department of Commerce

Donald L. Evans, Secretary

Technology Administration

Phillip J. Bond, Under Secretary for Technology

National Institute of Standards and Technology

Arden L. Bement, Jr., Director
Disclaimer

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

The policy of NIST is to use the Intemational 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.

Use in Legal Proceedings

No part of any report resulting from 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 Special Publication 1000-5

Natl. Inst. Stand. Technol. Spec. Publ. 1000-5, 1,054 pages (June 2004)

CODEN: NSPUE2

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 2004

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
Master Table of Contents

Volume 1

Table of Contents

List of Acronyms and Abbreviations

IVIetric Conversion Table

Preface

Executive Summary

Chapter 1 Interim Findings and Accomplishments

Chapter 2

Progress on the World Trade Center Investigation

Chapter 3

Update on Safety of Threatened Buildings (WTC R&D) Program

Chapter 4

Update on WTC Dissemination and Technical Assistance Program

Volume 2

Table of Contents

List of Acronyms and Abbreviations

Metric Conversion Table

Appendix A

Interim Report on the Analysis of Building and Fire Codes and Practices

Appendix B

Interim Report on Development of Structural Databases and Reference Models

for the WTC Towers

iii
Table of Contents

Appendix C

Interim Report on Analysis of Aircraft Impact into the WTC Towers

Volume 3

Table of Contents

List of Acronyms and Abbreviations

Metric Conversion Table

Appendix D

Interim Report on Preliminary Stability Analysis of the WTC Towers

Appendix E

Interim Report on Contemporaneous Structural Steel Specifications

Appendix F

Interim Report on Inventory and Identification of Steels Recovered from the

WTC Buildings

Volume 4

Table of Contents

List of Acronyms and Abbreviations

Metric Conversion Table

Appendix G

Interim Report on Significant Fires in WTC 1, 2, and 7 Prior to September 11, 2001

Appendix H

Interim Report on Evolution of WTC Fires, Smoke, and Damage based on Image Analysis

Appendix I Interim Report on Assessment of Sprayed Fireproofing in the WTC

Towers—Methodology
June 2004 Progress Report

Volume 5

Table of Contents

List of Acronyms and AbbreviatioQS

Metric Conversion Table

Appendix J

Interim Report on Experiments to Support Fire Dynamics and Thermal

Response Modeling

Appendix K

Interim Report on Subsystem Structural Analysis of the WTC Towers

Appendix L

Interim Report on WTC 7

Appendix M

Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire

Volume 6

Table of Contents

List of Acronyms and Abbreviations

Metric Conversion Table

Appendix N

Interim Report on Analysis of First-Person Accounts from Survivors of the WTC

Evacuation on September 11, 2001

Appendix O

Interim Report on Telephone Interviews

Appendix P

Interim Report on Emergency Communications

Appendix Q

NIST's Working Hypothesis for Collapse of the WTC Towers
Table of Contents

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vi
List of Acronyms and Abbreviations

AAPOR American Association of Public Opinion Research

ABC American Broadcasting Company

ACI American Concrete Institute AISC American Institute of Steel Construction

AISI American Iron and Steel Institute ALE Arbitrary-Lagrangian-Evlerian

AMCBO Association of Major City/County Building Officials

ANSI American National Standards Institute ANSYS finite element model

ARA Applied Research Associates, Inc. ASCE American Society of Civil Engineers

ASHRAE American Society of Heating, Refrigerating and Air-Conditioning Engineers, Inc. ASME American Society of Mechanical Engineers

ASTM ASTM International

AWS American Welding Society

BOCA Building Officials and Code Administrators

BOCA/BBC BOCA Basic Building Code

BPAT Building Performance Assessment Team

BPS Building Performance Study

BSI British Standards Institution

C/F cancer free P ATT cuiiipuicr-dsaisicu iciepiiuuc iiiicivicwa CBR chemical, biological, and radiological

CBS Columbia Broadcasting System

CERF Civil Engineering Research Foundation

CFD computational fluid dynamics

CIB International Council for Research and Innovation in Building and Construction

CII Construction Industry Institute CNN Cable News Network

vii
List of Acronyms and Abbreviations

CPP Cermak Peterka Peterson, Inc. CPU central processing unit

CRT cathode-ray tube

CTB&UH Council on Tall Buildings and Urban Habitat

CTE coefficients of thermal expansion

DC/F BlazeShield DC/F fire protective insulation

DL dead load

DTAP dissemination and technical assistance program

EMS Emergency Medical Service

EMT Emergency Medical Team

ER&S Emory Roth & Sons

FBI Federal Bureau of Investigation

FCA Flux cored arc FDNY New York City Fire Department

FDS Fire Dynamics Simulator

FE finite element

FEA finite element analysis

FEM finite element model

FEMA Federal Emergency Management Agency

FMRC Factory Mutual Research Corp.

FSI Fire-Structure Interface

FVM Finite Volume Method

GFI Government Furnished Information

GG glass over glass

GHz gigahertz

GMS, LLP Gilsanz Murray Steficek, LLP

HAZ heat affected zone HNSE Hugo Nue Schnutzer East HRR heat release rate HVAC heating, ventilating, and air conditioning

lAQ indoor air quality

IBC International Building Code

viii
June 2004 Progress Report

ICBO International Conference of Building Officials

ICC International Code Council

IMTI Integrated Manufacturing Technology

JFK John F. Kennedy International Airport

SIS Japan Industrial Standard

LERA Leslie E. Robertson Associates

LES Large Eddy Simulation

LL live load

LSTC Livermore Software Technology Corporation

MBC BOCA National Building Code

MCC Municipal Code of Chicago

MPI Message Passing Interface NBC National Broadcasting Company

NBFU National Board of Fire Underwriters

NCSBCS National Conference of States on Building Codes & Standards, Inc NCST National Construction Safety Team

NEMA National Electrical Manufacturers Association

NFPA National Fire Protection Association

NIBS National Institute of Building Sciences

NIST National Institute of Standards and Technology

NYC New York City

NYCBC New York City Building Code

NYCDOB New York City Department of Buildings

NYPD New York City Police Department

NYSBC New York State Building Construction Code

P.L. Public Law

PANYNJ Port Authority of New York and New Jersey

PAPD Port Authority Police Department

PC&F Pacific Car and Foundry

PDM Pittsburg-Des Moines

PONYA Port of New York Authority

R«&D research and development
List of Acronyms and Abbreviations

RWDI Rowan Williams Davis and Irwin, Inc. SBCCI Southern Standard Building Code

SDL superimposed dead load

SDO standards development organization

SEAoNY Structural Engineers Association of New York

SFPE Society of Fire Protection Engineering

SFRM spray-on fire resistant material or sprayed fire resistive materials

SHCR Skilling, Helle, Christiansen, & Robertson

SI metric . SLB short legs back-to-back

SMA Shielded Metal Arc

SOD Special Operations Division

SOM Skidmore, Ownings & Merrill

SPH Smoothed Particle Hydrodynamics

SQL Structured Query Language

SWMB Skilling, Ward, Magnussen, and Barkshire

TL Truss Lower Chord

TM Truss Middle Chord

TU Truss Upper Chord

UBC Uniform Building Code

UL Underwriters' Laboratories, Inc. use United States Code

USM United States Mineral Products Co.

VCBT Virtual Cybernetic Building Testbed

WABC WABC-TV New York

WCBS WCBS-TV New York

WF wide flange (a type of structural steel shape now usually called a W-shape). ASTM A 6

defines them as "doubly-symmetric, wide-flange shapes with inside flange surfaces that are substantially parallel."

WNBC NBC4 New York

WNYW F0X5 New York

WPIX WPIX-TV New York ^-^ WTC World Trade Center
June 2004 Progress Report

WTC 1 World Trade Center Tower 1 WTC 2 World Trade Center Tower 2 WTC 7 World Trade Center Building 7

XI
List of Acronyms and Abbreviations

Abbreviations

X by + plus or minus

°C degrees Celsius

°F degrees Fahrenheit

micrometer

2D two dimensional

3D three dimensional

cm centimeter

ft foot

ft^ square foot

yield strength (AISC usage)

g acceleration (gravity)

g gram

gal gallon

h hour

in. inch

kg kilogram

kip a stress unit equal to 1 ,000 pounds

kJ kilojoule

kN kilonewton

kPa kilopascal

klb 1 ,000 pounds

ksi 1 ,000 pounds per square inch

kW kilowatt

kW/m- kilowatts per square meter

L liter

lb pound

m meter

m' square meter

mm millimeter _

m/s meters per second
June 2004 Progress Report

min minute

MJ megajoule

MPa megapascal

mph miles per hour

ms microsecond

Msi millions pounds per square inch

MW megawatt

N newton

Pa pascal

DCf pounds per cubic foot

plf pounds per linear foot

psf pounds per square foot

psi pounds per square inch

s second

XUl
List of Acronyms and Abbreviations

This page intentionally left blank.

xiv
Metric Conversion Table

To convert from to Multiply by

AREA AND SECOND 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

6.4516 E-04

6.4516 E+00

8.361 274 E-01

ENERGY (includes WORK)

kilowatt hour(kW* h)

quad (1015 BtulT)

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

3.6E+06

1.055 056 E+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)

kipd kip=1000 Ibf)

kip(l kip=1000 Ibf)

pound-force (Ibf)

newton (N)

newton (N)

newton (N)

newton (N)

kilonewton (kN)

newton (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 DIVIDED 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 RATE

calorieth per minute (calth/min)

calorieth per second (calth/s)

kilocalorieth per minute (kcalth/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

XV
Metric Conversion Table

To convert from to Multiply by

LENGTH

foot (ft) meter (m) 3.048 E-01

inch (in) meter (m) 2.54 E-02

inch (in) centimeter (cm) - 2.54 E+00

micron (m) meter (m) l.OE-06

yard (yd) meter (m) 9.144 E-01

MASS and MOMENT OF INERTIA

kilogram-force second squared

per meter (kgf * s7m) kilogram (kg) 9.806 65 E-HOO

pound foot squared (lb * ft^) kilogram meter squared (kg * m^) 4.214 011 E-02

pound inch squared (lb * in'^) kilogram meter squared (kg * m") 2.926 397 E-04

ton, metric (t) kilogram (kg) l.OE+03

ton, short (2000 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

(not 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/ft") 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

xvi
June 2004 Progress Report

To convert from to Multiply by

TEMPERATURE

degree Celsius (°C)

degree centigrade

degree Fahrenheit (°F)

degree Fahrenheit (°F)

kelvin (K)

kelvin (K)

degree Celsius (°C)

degree Celsius (°C)

kelvin (K)

degree Celsius (°C)

T/K = t/°C + 273.15

t/ °C « t /deg. cent.

t/°C = (t/°F2 32)/1.8 T/K = (t/°F + 459.67)/l.

t/°C = T/K2 273.15

TEMPERATURE INTERVAL

degree Celsius (°C)

degree centigrade

degree Fahrenheit (°F)

degree Fahrenheit (°F)

degree Rankine (°R)

kelvin (K)

degree Celsius (°C)

degree Celsius (°C)

kelvin (K)

kelvin (K)

1.0 E+00

l.OE+00

5.555 556 E-01

5.555 556 E-01

5.555 556 E-01

VELOCITY (includes SPEED)

foot per second (ft/s)

inch per second (in/s)

kilometer per hour (km/h)

mile per hour (mi/h)

mile per minute (mi/min)

meter per second (m/s)

meter per second (in/s)

meter per second (m/s)

kilometer per hour (km/h)

meter per second (m/s)

3.048 E-01

2.54 E-02

2.777 778 E-01

1.609 344 E+00

2.682 24 E+01

VOLUME (includes CAPACITY)

cubic foot (ft^)

cubic inch (in^

)

cubic yard (yd^)

gallon (U.S.) (gal)

gallon (U.S.) (gal)

liter (L)

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

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

cubic meter (m')

cubic meter (m^)

cubic meter (m^)

cubic meter (m^)

liter (L)

cubic meter (m^)

cubic meter (m^)

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

xvii
Metric Conversion Table

This page intentionally left blank.

xviii
Table of Contents

List of Figures J-ii

List of Tables J-i

Appendix J Interim Report on Experiments to Support Fire Dynamics and Thermal

Response Modeling J-1

J.l Introduction J-1

J. 2.1 Description of Experiments J-2

J. 2. 2 Preliminary Results J-7

J. 3 Experiments for Guiding the FDS Fire Growth Predictions J-1

2

J. 3.1 Description of the Experiments J-12

J . 3 . 2 Preliminary Results J- 1

8

J. 4 Experiments for Fire Model Validation J-27

J.4.1 Preliminary Results J-32

J. 5 References J-34

Attachment 1 Simulating the Fires in the World Trade Center J-35

J-i
List of Figures

Figure J-1. Experimental enclosure during construction, viewed with access panels removed J-3

Figure J-2. Compartment content layout

Figure J-3. Tubular column J-5

Figure J^. Bar joist J-5

Figure J-5. Simple bar J-6

Figure J-6. Measured heat release rate as a function of time during tests 1-6 J-7

Figure J-7. Temperature-time history for truss A in test 5 J-8

Figure J-8. Temperature-time history for truss A in test 3 J-9

Figure J-9. Comparison of gas temperatures, test 3, exhaust side of compartment J-10

Figure J-10. Comparison of gas temperatures, test 5, exhaust side of compartment J-10

Figure J-1 1 . Comparison of heat fluxes to column, test 3 J-1

Figure J-12. Comparison of heat fluxes to column, test 5 J-1

Figure J-1 3. Photograph of right side of generic workstation J-1

3

Figure J- 14. Photograph of left side of generic workstation J-1

3

Figure J-15. Photographs of high-end workstation J-15

Figure J-1 6. Photograph of workstation soon after ignition in test 2 J- 18

Figure J-1 7. Photograph of workstation near the peak heat release rate in test 2 J-1

9

Figure J 18. Heat release rate versus time for single workstation test 1 J-19

Figure J-19. Heat release rate versus time for single workstation test 2 J-20

Figure J-20. Heat release rate versus time for single workstation test 3 J-20

Figure J-21. Heat release rate versus time for single workstation test 4 J-21

Figure J-22. Heat release rate versus time for single workstation test 5 J-21

Figure J-2 3. Heat release rate versus time for single workstation test 6 J-22

Figure J-24. Preliminary comparison of predicted and measured HRR values J-26

Figure J-25. Plan view of test configuration J-28

Figure J-26. Elevation view of test configuration J-29

Figure J-27. Views of interior of test fixture J-30

Figure J-28. View of fire compartment before the start of test 6 J-33

Figure J-29. View of fire compartment 2 minutes after the start of test 6 J-33

Figure J-30. Fire Dynamics Simulator 3D rendition of experimental enclosure J-34

Figure J-3 1 . Comparison of measured and predicted heat release rate for test 1 J-34

J-ii
List of Tables

Table J-1 . Test matrix J-6

Table J-2. Summary of insulation on steel components J-6

Table J-3. Peak heat release rates and times to peak J-22

Table J^. Test matrix J-3

1

J-i
Appendix J

This page intentionally left blank.

J-iv
Appendix J

Interim Report on Experiments to Support Fire Dynamics and

Thermal Response Modeling

J.I introduction

The reconstruction of the World Trade Center (WTC) fires involves two computation models:

• Fire Dynamics Simulator (FDS): This is the first large-domain 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 model numerically solves the conservation equations of

mass, momentum and energy that govern low-speed, thermally driven flows with an emphasis on smoke and heat transport from fires. The companion software package, called Smokeview,

graphically presents the results of the FDS three-dimensional (3D) time-dependent simulation.

Smokeview animates in three dimensions the smoke particulates, heat fluxes, temperatures and

fluid velocities within a building. Users of the package can view the enclosure from any angle

and from inside or outside.

• Fire-Structure Interface (FSI): This code effects 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 members, such as ANSYS. The members may be simple (e.g., bare steel) or complex (e.g., insulation-coated steel).

For application to the Investigation, each of these 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 will be

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

The following text describes three sets of experiments designed to accomplish this. All three sets have

been completed, and the analysis of the data is under way. Full reports are in preparation and will be

completed in summer 2004. Attachment 1 is a short paper on the modeling and experiments that will be

presented at Interflam in July 2004.

J.2 experiments for accuracy assessment of thermal

environment modeling

The purposes of these experiments, conducted in the National Institute of Standards and Technology

(NIST) Large-scale Fire Laboratory from March 10 through 26, 2003, were to:

• Assess the accuracy with which FDS predicts the thermal environment in a burning compartment

and;

J-1
Appendix J

• Establish a data set to validate the prediction of the temperature rise of structural steel elements

using FSI.

Within a large test compartment, assorted steel members were exposed to controlled fires of varying heat

release rate (HRR) and radiative intensity. The steel members were bare or coated with spray-applied

fireproofmg of two thicknesses. The thermal profile of the fire was measured at multiple locations within

the compartment. Temperatures were also recorded at multiple locations on the surfaces of the steel, the

insulation, and the compartment. Prior to each test, a prediction of the thermal envirormient in the compartment was determined using FDS. Following the tests, the prediction and experimental results were compared.

J.2.1 Description of Experiments

The test compartment consisted of a steel stud frame lined with calcium silicate board. The internal dimensions of the compartment were 3 m high, 7 m deep, and 4 m wide. There were four openings in the west wall through which air entered the room; they totaled 1.75 m" (10.8 ft") in area and were located 1 m

(3.3 ft) above the floor. There were four openings in the east wall through which heat and combustion

products were emitted; they also totaled 1.75 m" (10.8 ft") in area and were located 2 m above the floor. A photograph is shown in Fig. J-1 and a schematic in Fig. J-2.

In each of the six tests, the four test subjects were a bar, two trusses, and a thin-walled tubular column.

These are depicted in Figs. J-3 through J-5. Depending on the test, these specimens were either left unprotected or were coated with spray-applied fire protective insulation material, BlazeShield DC/F. The

fibrous insulation was applied by an experienced applicator who took considerable care to apply an even

coating of the specified thickness. As such, the insulated test subjects represent a best case in terms of

thickness and uniformity.

The fires consisted of liquid hydrocarbon fuels sprayed by a two-nozzle spray burner onto a 1 m x 2 m

(3.3 ft X 6.6 ft) pan. The fuels were (a) heptanes and (b) a mixture of nominally 60 percent (by mass)

heptanes with 40 percent toluene. The latter fuel produced a significantly sootier flame.

The instrumentation for the tests comprised up to 352 channels of data.

• The combustion products were collected in a 6 m x 6 m (21.5 ft x 21.5 ft) hood. Instrumentation

in the exhaust duct enabled calculation of the rate of heat release throughout a test.

• Fourteen heat flux gauges were placed strategically around the compartment to measure the

transport of radiant energy; in addition, there were four slug calorimeters measuring the total heat

flux parallel to the trusses.

Most of the channels were for thermocouples that measured the temperatures on the surface of the walls

and ceiling, within the walls, on the surface of the steel components, and at the surface of the spray- applied insulation. With the large number of measurements, it was possible to go beyond the traditional point-by-point comparison and discover why the model either under- or overpredicted a given

measurement. A description of the test series appears in Table J 1. Table J-2 shows the dimensions and

J-2
Interim Report on Experiments to Support Fire Dynamics

variability of the spray-applied insulation. The measurements were taken at numerous locations along the perimeter and length of each specimen using a pin-thickness gauge specifically designed for this type of

insulation.

Figure J-1. Experimental enclosure during construction, viewed with access

panels removed.

J-3
Appendix J

PLAN VIEW

7.04 m NORTH

Baffle Column

Trusses

3.68 m

Inlets WEST

(0,0 Y

X

Inlet Vents

Fire Pan

Susoended bars

Fire Pan

Access panels

1.2 SECTION A-A

3.68 m

Trusses

Truss support

Baffle •

(0,0)

Truss support

iO Outlets

-Suspended

bar

3.82 m

NORTH

Column

Column

support

Figure J-2. Compartment content layout (not to scale).

J-4
Appendix J

A—y\

3.35 m

^.31 mm

6.25 mm

T

0.26 m

\

^ 0.36 m

*

Figure J-3. Tubular column.

4.6 m !

1 1

0.8 m \ ^^^^^^^^

Web Bar Top Chord Angles

-^c r A /I 76 mm

—*\J^— 25.6 mm 6.4 mm

^ ^

6.5 mm

Bottom Chord Angles

6.4 mm —n* n

76 mm 6.3 mm

I

64 mm

<

64 mm

Figure J-4. Bar joist.

J-5
Appendix J

25.5 mm I 3 m length

Figure J-5. Simple bar.

Table J-1 . Test matrix.

Test Measured HeatRelease Rate (MW) Fuel

Planned Insulation

Thickness (mm)

Planned Test Duration (min)

1 2.0 Heptanes None 15

2 2.4 Heptanes/toluene None 15

3 2.0 Heptanes/toluene None 15 4 3.2 Heptanes Same as test 5 15

5 3.0 Heptanes See Table J-2 50

6 3.0 Heptanes See Table J-2 50

Table J-2. Summary of insulation on steel components.

Test Item Specified Thickness (mm)

Applied Thickness (mm)

Mean Std. Deviation

5 Bar 19.1 23.0 5.5 Column 38.1 41.0 3.0

Truss A 19.1 26.9 7.3

Truss B 38.1 40.5 8.2

6 Bar 19.1 25.3 4.6 Column 19.1 21.4 3.5

Truss A 19.1 26.0 6.9

Truss B 19.1 25.6 6.9

The HRR are shown in Fig. J-6. The important features of the HRR curves are:

• The expected (from the calorific value of the fuels) and measured heat release rates agreed

within the relative expanded experimental uncertainty of + 1 1%.

• The heat release rates are steady over the time interval in which the burner was turned on.

J-6
Interim Report on Experiments to Support Fire Dynamics

3500

-e— Test 1 -a— Test 2

-A— Test 3

O " Test 4 •— Test 5

Teste

1500

1000

0 M.

0 1000 2000

Time (s)

3000 4000

Figure J-6. Measured heat release rate as a function of time during tests 1-6.

In each test, the baseline signals from all the measurement devices were established; then the burner was

ignited and continued burning at a steady rate until the temperature at any steel surface approached

approximately 600 °C. (Above this temperature, there was concern that loss of strength might lead to collapse and accordant damage to the test facility.) At that point, the burner was turned off. In test 2, the

steel reached the target temperature at about 6 min and was terminated at that time. During test 4, the 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. Thus the test was terminated early, and the data

have not been processed. 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 thennocouple extensions were

visually inspected and were found to be undamaged.

J.2.2 Preliminary Results

Figure J-7 shows typical temperature data obtained in the tests. These data are for truss A in test 5. The

thermocouple location notation is as follows: TU: Truss Upper Chord, TM: Truss Middle (Web), TL:

Truss Lower Chord; 1 to 4: locations across the length of the test specimen; S: on the steel surface, I: on

the outer surface of the insulation; A: truss A. For the informative nature of this progress report, it is only

important to note the following:

• The curves that rise sharply from the beginning of the test are those for temperatures on the outside of the insulation.

• The curves that rise more gently are those for temperatures at the interface between the steel and

the insulation.

J-7
Appendix J

From the figure, one can see that:

• The 19. 1 mm (0.75 in.) insulation delays the rise to a peak steel temperature by almost an hour at

all locations.

• The highest temperature reached at the steel surface is approximately 300 °C lower than the

temperature at the outside face of the insulation material.

The curve patterns for the other steel specimens in the tests with insulated steel are similar in shape.

Test S: Truss A

600 1200 1800 2400 3000

Time (sec)

3600 4200 4800

-TUISA

-TU2SA TU3SA

-TU4SA -TMISA

-TM2SA

-TM4SA

-TLISA TL2SA

TL3SA

TL4SA

TUIIA

-TU2IA

-TU3IA TU4IA -TM2IA

-TM3IA TM4IA

TLIIA

-TL2IA

-TL3IA TL4IA

Figure J-7. Temperature-time history for truss A in test 5.

By contrast. Fig. J 8 shows the same plot for truss A from test 3, in which the truss was not insulated and

the fire was of shorter duration and lower intensity. Nonetheless, the outer surface of the steel reached the targeted maximum temperature (just short of 600 °C) in about one-third the time. As expected, this result

is typical of the fire response of the uninsulated steel specimens in tests 1 through 3.

Figures J-9 through J- 12 show comparisons of the modeled and measured temperatures and heat fluxes

for tests 3 (2.0 MW heptanes/toluene fire) and test 5 (3.0 MW heptanes fire). The agreement for the

highest temperatures is excellent. Analysis of both the model and the thermocouple measurements is under way to determine the source of their differences at the lower temperatures, especially in test 5. The

spikes in the heat flux plots are artifacts—the result of the periodic nitrogen blasts to reduce soot accumulation on the gage surface.

J-8
Interim Report on Experiments to Support Fire Dynamics

Figure J-8. Temperature-time liistory for truss A in test 3.

While the data analysis is not yet complete, the following are valuable preliminary observations:

The prediction of the upper layer temperatures was within experimental uncertainty. 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 prediction of the time for the steel surfaces

to reach 600 °C was accurate.

The sootier burning fuel led to similar temperature rise in the ceiling and the steel above the fire plume.

• The model predicted the asymmetric shape of the fire plume, caused by obstructions to uniform

flow through the compartment.

Further analysis will use Smokeview for visual comparison of the test results and the model predictions.

This will determine how well FDS captures both the fire phenomena and the thermal patterns in the compartment. Quantitative analysis of the data will then determine the numerical accuracy of the

predictions. Similar analysis will be performed to determine the accuracy of the finite element modeling

of the thermal patterns within the bare and insulated steel components.

J-9
Appendix J

U

600

500 -

400

"i 300

B

H 200 -

100

WTC Phase 1, Test 3, East Aspirated TCs

Figure J-9. Comparison of gas temperatures, test 3,

exhaust side of compartment.

1200

1000

u

B

(UB

<L) H

800 -

600 -

400 -

200 -

WTC Phase 1, Test 5, East Aspirated TCs

365 cm (Exp)

215 cm (Exp)

35 cm (Exp)

365 cm(FDS)

215 cm(FDS)

35 cm (FDS)

1000 2000

Time (s)

3000

Figure J-10. Comparison of gas temperatures, test 5,

exhaust side of compartment.

j-io
Interim Report on Experiments to Support Fire Dynamics

1000

Time (s)

Figure J-1 1 . Comparison of heat fluxes to column, test 3.

1200

S

5

X

100

80 -

60 -

Z 40 -

C3

20 -

High North (Exp)

Low North (Exp)

High North (FDS)

Low North (FDS)

WTC Phase 1, Test 5, Column Heat Flux Gauges, North Facing

1000 2000

Time (s)

3000

Figure J-12. Comparison of heat fluxes to column, test 5.

J-ll
Appendix J

J.3 EXPERIMENTS FOR GUIDING THE FDS FIRE GROWTH PREDICTIONS

In the early stages of this Investigation, the FDS combustion module was enhanced to enable the

inclusion of charring materials, such as those that comprise much of the office furniture. Thermophysical

property data from the combustion of small (100 mm x 100 mm) samples of the furnishings were

obtained using the Cone Calorimeter. These data became input to the fire simulation. A set of real-scale experiments was then designed and performed (July and August 2003) to identify any need for further enhancements to the fire model. In each of these experiments, a single workstation (i.e., an office cubical

or module), similar to those in the WTC offices, was burned under a hood. A soffitted ceiling allowed for the collection of hot fire effluent and the accordant thermal radiation to the test specimen. Some of the

tests examined the effects on the burning rate ofjet fuel and/or noncombustible material occluding a

fraction of the workstation surfaces.

J.3.1 Description of the Experiments

Materials

Workstations come in a variety of styles and finishes. However, they tend toward similar size and mass. They are also fabricated of materials with similar burning behavior, e.g., the work surfaces are generally

laminated particleboard or wood. Most of the workstations burned here were of a single generic type.

For comparison, one high-end unit (identical to those in the aircraft impact floors in WTC 1) was also

burned in a manner identical to one of the tests of the generic units.

The generic workstation examined is shown photographically in Figs. J-13 and J-14. The layout,

including the placement of the various nonstationary items, was suggested by personnel from a company

that supplied office furnishings to the occupants of WTC 1. Information on the distribution of papers and

other office items was provided by a frequent visitor to these offices. The workstation covered a footprint

nominally 2.44 m x 2.44 m (8 ft x 8 ft) and was surrounded by privacy panels.

• The panels were 1 .22 m (4 ft) high, except that on one side the panel was 1 .52 m (5 ft) tall and

supported a bookcase. On the side opposite the bookcase, there was a 1 .22 m (4 ft) wide open

entrance opening. The panels were made of a steel and softwood frame, covered on both sides

with layers of fiberglass padding and perforated steel and a thermoplastic cover fabric. A few

sheets of copier paper were tacked to the cubicle walls on three sides.

• The work surfaces were formed from four sections of laminated medium density fiberboard

supported by steel brackets from the wall panels. Four document boxes contained a total of four reams of copier paper. Additional paper was stacked horizontally on the desk surface.

• The seat and back of the office chair were a nonthermoplastic fabric over polyurethane foam

supported by a one-piece thermoplastic shell; its five-legged base was thermoplastic with steel framing and support elements.

J-12
Interim Report on Experiments to Support Fire Dynamics

Figure J-13. Photograph of right side of generic workstation.

Figure J-14. Photograph of left side of generic workstation.

J-13
Appendix J

• The three file cabinets (0.91 m wide, 0.51 m deep, 0.68 m high [36 in. x 20 in. x 27 in.], with two

horizontal drawers) were painted steel; they rested directly on the carpet. Two of the cabinets

contained two reams each of copier paper as a rough means of assessing the extent to which paper

in file drawers might contribute to a fire.

• The bookcase (1 .22 m [48 in.] long) had a steel shelf and top but these were supported only on

their ends by combustible end panels; the steel front closure panel was fabric-covered steel and it was open (on top of the bookcase). Ten document boxes held about 13 reams of copier paper.

• The carpet tiles were nylon fiber-faced over a dense foam rubber backing. A square area 2.74 m X 2.74 m (9 ft x 9 ft) was covered with 36 carpet tiles.

• The computer monitor was a nominally 1 7 in. CRT-based unit. Its front face was taped with

fiberglass tape and it was pointed toward the wall panel opposite the cubicle opening for safety in the event of an implosion. The keyboard was placed in its normal location, parallel to the sloped

segment of the work surface. The computer processor (tower-type container with plastic only on

the firont face of the container) was placed on the floor next to a waste paper basket (both on the

side opposite the cubicle opening).

• The wastebasket was thermoplastic and contained one ream of copier paper atop five balled-up

paper ream wrappers.

Thermophysical characterizations of six of the generic workstation materials (carpet, panels, work

surface, chair seat, paper stack, and computer monitor shell) were obtained using the Cone Calorimeter.

These data were to serve as input to FDS. Since the physical behavior of at least some of the work station

materials (and the objects from which they were taken) was expected to be more complex than was

revealed in the Cone tests, it was anticipated that these full-scale tests would provide clues as to necessary

empirical adjustments in the FDS predictions.

The high-end unit was similar to the generic workstation, with the principal differences being:

• The wall panel construction was somewhat different having a 3 mm (0.125 in) layer of flameretarded polyester fiber beneath the outer fabric, a more open steel panel beneath this, a central fiberboard layer (3 mm thick) and an all steel peripheral frame (no wood). The fiberboard

roughly doubled the amount of woody fuel within the wall panels of the enclosure and put it into a much higher surface area form in which it could be expected to bum appreciably faster. It should be noted, however, that this increase in woody fuel was only about 10 percent to 15 percent of the total available in the desk surfaces. Also, its enclosure deep within the wall

panels delays its burning.

• The file cabinets (four, with a total face length of 2.67 m rather than three with a total face length

of 2.44 m) had a flammable, charring plastic surface on the drawer fronts that added 15 percent to 20 percent of flammable area.

The chair was constructed somewhat differently (seat and back as separate pieces) and behaved as if its upholstered surfaces were flame-retarded. The high-end workstation is shown in Fig. J-15.

J-14
Interim Report on Experiments to Support Fire Dynamics

Figure J-15. Photographs of high-end workstation.

Test Configuration and Instrumentation

The workstation to be tested was placed on top of a double layer of 13 mm ( 1/2 in) thick calcium silicate sheets. These in turn rested on a set of four weighing cells, one at each comer. Each workstation

(including furnishings) weighed approximately 730 kg (1,600 lb) and contained approximately 300 kg

(660 lb) of combustible material. The weighing cells are accurate to ±0.1 kg (0.2 lb). The entire assembly was placed beneath the hood of the NIST 10 MW calorimeter hood to allow continuous heat

release rate measurement.

J-15
Appendix J

A fire typically forms a layer of hot, smoky combustion products near the ceiling of a room. Thermal

radiation from this layer plays a significant role in fire spread. Thus the test fixture included partitions to hold the combustion products from these test fires. The ceiling was a 3.66 m x 3.66 m (12 ft x 12 ft) section of 13 mm thick calcium silicate board. It was supported on a water-cooled steel frame 2.74 m

(9 ft) above the floor of the workstation. To keep the gases from flowing quickly across the ceiling and

thus not forming an appropriate layer, the ceiling was surrounded on all four sides by a steel skirt that draped down 0.61 m (2 ft) from the ceiling.

The test instrumentation included:

• Instrumentation in the hood exhaust duct for measurement of rate of heat release.

• Four video cameras placed to record the progression of flame spread over the objects in the

cubicle. Their view of the combustibles became observed by the wall panels and the flames

themselves as the fire grew in intensity. An observer narrated the fire growth to supplement what

the cameras recorded directly.

• In two tests, an upward-facing, water-cooled Schmidt-Boelter total heat flux gage 127 mm (5 in)

above the cubicle floor, near the cubicle center.

• An external total heat flux gage mounted so as to view the entire fire from one side. This

provided a signal that was proportional to the instantaneous HRR of the fire and was useful for certain timing issues.

• In two tests, six 24-gage chromel/alumel thermocouples to follow the progress of the fire on the

underside of the desk surfaces.

Ignition Scenario

The initiation of the tests was in keeping with the workstation being one of a large array on a given floor of a large office space and a fire propagating through the array. Thus a large ignition source (a 2 MW,

four-nozzle spray burner over a 2 m x 1 m pan) was placed immediately adjacent to the exterior of one

wall panel of the test station, simulating the burning of the adjacent workstation. The size and placement

(pan bottom 0.81 m above the floor) of this ignition source were guided by preliminary FDS predictions.

The igniter fire was supported by a flow of commercial-grade liquid heptanes (mixed isomers) sprayed at a nominally steady rate from the four nozzles pointing downward toward the pan. The heptane mix was

supplied from a reservoir by a variable-speed pump. The desired fuel flow was preset before the test and

measured in triplicate by catching the flow from each nozzle in a volumetric cylinder for a typical period

of 20 s. The nominally 2 MW fire supported by the heptane flow typically impinged almost continually on the ceiling above the igniter. There was thus an essentially continuous wall of flames radiating toward

the workstation along the central three-fourths of the length of one panel. In addition, the workstation was subject to radiation from the hot ceiling and the hot smoke captured below the ceiling.

J-16
Interim Report on Experiments to Support Fire Dynamics

Test Variables

There are numerous variables that might influence the burning of a workstation such as that examined

here, starting with the nature of the materials and their spatial arrangement, the ignition conditions, etc. There are also variables unique to the WTC fires: the amount of paper-based clutter in a workstation, the

possible presence ofjet fuel (and its amount), the possible presence and amount of inert rubble generated

by the airplane impact (fallen ceiling tiles, inert dust from pulverized concrete and/or wallboard), varying

degrees of impact-induced break-up and compaction of the work station itself, and presence of

combustible solids from the airplane.

The tests were designed keeping the purposes in mind: (a) identify the current capability of FDS to predict complex burning behavior using combustion data from small-scale specimens and (b) obtain clues

to improving the combustion algorithm should the predictions be of insufficient accuracy. Thus the two

selected variables were those that could affect the workstation fires in manners that test FDS and are important in the WTC context: the presence or absence of both jet fuel and of inert rubble. The inert rubble was taken to be representative of fallen ceiling tiles.

The primary test set thus focused on two levels of two variables examined in a full factorial design. This

calls for four tests: (1) no inert rubble and no jet fuel, (2) rubble but no jet fuel, (3) jet fuel but no rubble

and (4) rubble and jet fuel. The levels of the two variables were estimated to produce differences in burning behavior that would be clearly observable. Thus the inert rubble was chosen to cover approximately 30 percent of the horizontal surfaces facing the hot ceiling. As performed, 24 of the 40 ceiling tiles were on the horizontal desk surfaces; 14 were on the central, open floor area; and 2 were on the chair seat. The fraction of the plan view horizontal area covered by these tiles was approximately

31 percent. A total of 4 L (approximately 1 gal) of Jet A was spread over these same horizontal surfaces.'

Two additional tests were conducted. The first test (in the entire series) utilized what was nominally onehalf of a generic workstation, though it included both a full chair and full computer. This was done to gage the burning behavior of what was an entirely unknown system. The fourth test was of the high-end

workstation. Both of these were conducted with no jet fuel and no inert rubble present.

Test Procedure

The workstation was assembled a few hours before a test. Since the ambient humidity was high

(approximately 70 percent) for most of the tests, the paper was covered with plastic sheeting if it was to be exposed for more than 2 hours, although this probably did not preclude significant moisture pick-up in the outer portions of the paper piles. That moisture would be expected to somewhat slow the ignition

process relative to a more normal humidity of 50 percent.

At the beginning of each test, all instruments were calibrated. [The heat release rate was also measured

for a few minutes after the test ended in order to verify the calorimeter baseline.] The heptanes flow was

measured in triplicate. The test was initiated by starting the heptanes flow and immediately igniting it with a torch. This defined time zero. The heptanes flow was left on until late in the test unless there was

Areas close to the impact could have been drenched with higher quantities. The workstations would also have had to be

extensively fragmented, which is not a situation being examined here. Another pragmatic factor that kept the jet fuel loading

down was the probability that higher levels would have pushed the calorimeter beyond its maximum allowable capacity.

J-17
Appendix J

an indication that the fire was going to significantly exceed the 10 MW calorimeter capacity; this happened only once (test 5). The full test was video taped by all four cameras and a narration was fed to one camera, describing the sequence of ignition events that spread the fire over the accessible flammable

surfaces. In two cases (tests 3 and 4) the residual weight of the paper piles in the two file drawers was

obtained as a measure of the participation of this paper in the overall fire.

In the two tests in which jet fuel was placed on the horizontal surfaces prior to the start of a test, this was

done just before ignition of the 2 MW spray burner. A sprinkling can was used in two separate

operations, each involving 2 L of the liquid fuel. First, the liquid was sprinkled on the horizontal desk

surfaces and the objects on them (i.e., the various paper stacks or document boxes, the computer monitor

and the keyboard) using a timed movement that attempted to allot an equal amount of liquid to each onethird section of the total work surface. The desk surface had been leveled after installation so that the

liquid would not run preferentially in one direction. Next, an equal amount of the liquid was sprinkled in a similarly timed manner on the central open section of the carpet (not under the desk surfaces). Since the

chair occupied a portion of this space, the allotment for that portion went onto the chair seat and back

surfaces. For the objects on or in contact with the desk surface, there was some tendency for the jet fuel

to wick into them if they were porous. This was true of the paper, the inert tiles and the wall panel fabric

just above the desk surface.

J. 3.2 Preliminary Results

Figures J-I6 and J-17 show the workstation early in a test and at a time just before the peak heat release was reached. Figures J-18 through J-23 show the heat release rate curves from the six fire tests.

Figure J-16. Photograph of workstation soon after ignition

in test 2.

J-18
Interim Report on Experiments to Support Fire Dynamics

Figure J-1 7. Photograph of workstation near the peak heat

release rate in test 2.

6000 —i—I

1—

I

1 1 1

1 ?— —I 1

1 1 1 11—I 1—

I r—I

1 1—I

1 r

Time from Ignition (s)

Figure J-1 8. Heat release rate versus time for single workstation test 1.

J-19
Appendix J

1 10

8000

w 6000 -

a>

as DC

03

CO 1 4000

a:

CO

03 I

2000 -

1 1 1 1

'

' '

' 1 1 1 1 I 1 1 1

1

-

- k

-

1

l,J

!|= - O

V

03

! , . , , , , , , , , , , 1 1 1 1 1

0 500 1000 1500 ?000

Time from Ignition (s)

Figure J-19. Heat release rate versus time for single workstation test 2.

8000

7000

_ 6000

5000

4000

3000

2000

1000

0

E

cu

-t

I CO on

a>

CO

(O

q:

OJ X

.

' ' ' '

1

1 1 1 1 11 1 1 1 1 1 1 1 1 1 1 1

1 1 1 1 1

1^

0 I ; I c —

1 1 1 1 1 1 1 i 1 1 I 1 1 1 , , , , till 1 1 1 1 1

0 500 1000 1500 2000 2500 3000

Time from Ignition (s)

Figure J-20. Heat release rate versus time for single workstation test 3.

J-20
Interim Report on Experiments to Support Fire Dynamics

1 10

8000

§ 6000

a:

^ 4000

tr

05

2000

-T—I—I—rIgniter Off

1

'

v:

G ' '

i I I I I

0 500

-J—I I I I I I I I I I I I I I I I I

' 1^1 I' 1000 1500 2000 2500 3000

Time from Ignition (s)

Figure J-21. Heat release rate versus time for single workstation test 4.

1 10'

8000 -

fS 6000 -

CD

tr«

ai ^ 4000

2000 -

0

'—I—'—I' 500 1000 1500

Time from Ignition (s)

2000

Figure J-22. Heat release rate versus time for single workstation test 5.

J-21
Appendix J

1 10

8000

6000

^ 4000

2000 -

CD

a:

0)

(ii

CIS

-*

> CD

O) X

1 1 1 1

' '

' ' ' •

' 1 1 1 J

as

Falls

OutW

- \

\rul -mTW--

0)c

as D_^ O -

ID -t

1 1 1 1 1 . < 1

, , . , 1 1 1 r

1 1 1 1 500 1000 1500

Time from Ignition (s)

2000 2500

Figure J-23. Heat release rate versus time for single workstation test 6.

Table J-3 lists the peak HRR values from each test plus the time to that peak, rounded to the nearest 10 s. The HRR peak is given in two ways: first, as the absolute highest single reading recorded (for a 1 s interval) and second, as the average of the value at this peak plus values up to 5 s to either side of this peak. The latter compensates for both noise in the calorimeter system and in the fire itself; it is doubtful

that objects in a real room can respond in any meaningful way to small HRR fluctuations in a fire that is under 10 s in duration. The averaged values are about 3 percent lower than the absolute peak values, a

result of the sharpness of the peaks in every case.

Table J-3. Peak heat release rates and times to peak.

Test Test Specimen

Jet Fuel

(Yes/No)

Inert Tiles (Yes/No)Peak HRR

(kW)

Time to Peak

HRR (s)

1 Half of generic workstation N N 5920/5770 490

2 Generic work station N N 8700/8480 530

3 Generic work station N Y 7560/7300 590

5 Generic work station Y N 9120/8910 160

4 High-end work station N N 9890/9660 510

6 Generic work station Y Y 7960/7690 200

The videotapes of the fires show that the peak HRR corresponds closely to simultaneous burning of all the "accessible" combustible surfaces in the workstation interior. This included the top of the desk

surface, the objects on it, the exposed area of the materials in the bookcase, the full chair area (but see below), the exposed area of the carpet (i.e., not that under the three steel file cabinets), objects on the

J-22
Interim Report on Experiments to Support Fire Dynamics

carpet (computer processor case and wastebasket) and the underside of the desk, except above the steel

file cabinets where the air access was limited. The interior surfaces of the wall panels contributed

negligibly because the thermoplastic fabric melted and rolled downward into a mass having much less surface area than the original fabric prior to igniting. This type of behavior is not captured in Cone

Calorimeter tests.

The following observations emerge from comparison of the results in Table J 3

• The peak fire intensity from the half workstation is about two-thirds that of the frill workstation.

This is probably primarily the result of two factors:

- The same chair was present in both cases; this chair has an estimated HRR peak in the

neighborhood of 1/2 MW by itself.

- The inert steel file cabinets cover twice as much of the carpet in the second half of the workstation, precluding its participation in the HRR peak; this lowers the HRR

contribution from the second half of the workstation.

• The computer was also totally combusted within the first half of the experiments, though this was

a substantially lesser heat source than was the chair.

• The ceiling tiles reduced the peak HRR in proportion to their coverage of the burning surfaces; both just under 15 percent. While there was 30 percent coverage of the upward facing horizontal

surfaces, there was no coverage of the underside of the desk, the carpeting below the desk or the underside and back of the chair. As noted above, all of these were burning at the peak. (The

reduction might become more than linear if nearly all of the upward facing horizontal surfaces were covered since this could preclude the progressive flame spread that gets all accessible

surfaces involved)

• The principal effect of the presence of 4 L of Jet A on the horizontal surfaces was in shortening

the time to involvement of all accessible combustible surfaces, and thus the time to the peak

HRR. The peak itself was boosted upward only about 5 percent, presumably because the Jet A

helped boost the overall fuel gasification rate somewhat while adding its high heat of combustion.

When the inert tiles were also present, the Jet A was poured across their top surfaces temporarily

rendering these surfaces flammable. Since the tiles were porous, the Jet A burning rate on them was reduced, however, and the tiles still managed to produce a 13 percent to 14 percent reduction

in the peak HRR relative to the case with Jet A and no tiles.

From examination of the videotapes and the commentary, NIST determined that, when there was no Jet A

present, there was a fairly reproducible progression of ignition events leading up to the HRR peak:

• The onset of the igniter fire bathed the entire workstation in radiant heat.

• The igniter fire was not the pilot flame that ignited other objects, although FDS simulations

suggest that the peak fluxes on surfaces facing the igniter were of the order of 30 kW/m" or more,

well above the minimum flux for piloted ignition of the various exposed surfaces. Even the top

of the computer monitor shell, which gasified extensively only 20 cm from the spray burner

J-23
Appendix J

flames, did not appear to ever have directly been ignited by those flames. Instead the progression

of flaming was initiated by the top sheets on the paper stacks. Its presence in many places made

it highly instrumental in spreading the flames to the computer monitor shell, the desktop and the

top of the chair back. Flaming material from the chair fell onto the carpet, igniting it. Subsequent spread on the carpet was delayed, however, until the chair was fully involved, along

with the remainder of the upward-facing surfaces of the desktop. Later, paper ignition brought

flames to the two other sides of the desk area in the workstation.

• Once the entire area at and above the top of the desk surface was burning, the "compartment"

under the desk flashed over from the radiant heat from the upper compartment and especially the

heat from the burning chair. With all flammable surfaces ignited, the HRR quickly peaked. The

specific construction of the chair and its location were critical to this fire growth process:

- One comer of the seat was deliberately placed to extend approximately 15 cm under the

desk surface on which the computer keyboard rested. This assured that some heat would

reach this space early on, even though the chair flames would contact the relatively

ignition resistant underside of the desk.

- When only the upholstered surfaces of its seat and back were burning, the chair retained

its original shape, and little of its heat reached the lower compartment area.

- When the thermoplastic support shell of the chair began to melt and flow to the floor,

extensive heat flowed directly into the "compartment" under the desk.

- The partial steel skeleton kept the chair from collapsing and maintained burning from

desk level to floor level. The flame radiation to the cavity under the desk quickly ignited

the materials located there, while the upper part of the chair flames played on the desk,

assisting ignition of its underside.

Since chairs of different designs (and thus different burning behavior) could be fabricated from the same

materials, the detailed fire behavior of the chair cannot be inferred simply from the Cone Calorimeter data

for the component materials. Thus, empirical treatment of its HRR process is necessary.

As expected, the progression of ignition events in the presence of Jet A was different. Ignition of the materials in the cubicle occurred more quickly. However, the manner of the acceleration was not what

might have been expected.

• Since the flash point of Jet A is at least 46 °C (approximately 20 °C above the ambient

temperature), initially there would be no flammable mixture of vapors above the liquid fuel surface. As the strong radiant flux from the spray burner bathes the workstation, the videos show

an increasingly dense aerosol rising from the various wetted surfaces near the spray burner.

These did not ignite, as evidently turbulent mixing of the vapor plume above the surfaces quickly

diluted the fuel vapor below the lean flammable limit. Instead, in the first Jet A test, a random

piece of flaming material rose from the burner area and drifted down on top of the desk,

immediately igniting the Jet A. In the second test, a flaming piece of debris from the burner

landed on top of the bookcase where there was no jet fuel and it had no effect. The paper stacks

J-24
Interim Report on Experiments to Support Fire Dynamics

on the work surface next to the spray burner finally dried out, began to char, and then transitioned

into flaming, igniting the jet fuel.

• The Jet A flames then spread rapidly, but did not sweep continuously around the desktop.

Presumably the initial evaporation period left some dry spots that stopped the spread.

• The carpet was ignited by flaming matter dropping from the chair. Apparently the turbulent

vapor plume dilution process mentioned above prevented the flames from jumping downward to the carpet fi-om the flaming desk surface, even though the carpet was emitting an aerosol. At any

rate, the subsequent flame spread on the carpet was rapid.

• None of these steps is resolved in the HRR curves of Figs. J-21 and J-22 due to the rapidity with

which the curves jumped rapidly to their peak values soon after the Jet A ignited. As in the "dry"

tests, the peak corresponds to all accessible combustible surfaces burning simultaneously.

The rapid decay in HRR after the peaks in all tests presumably reflects several factors that should be

captured in the combustion algorithm in FDS:

• The various paper piles developed a thick ash layer that would drive down their burning rate.

• Char formation on the desk surfaces drove down its burning rate.

• The carpet began to bum out.

However, a number of geometric changes occurred that are beyond the capability of FDS to reproduce:

• The chair fire rapidly collapsed to a pool fire on the floor whose reduced burning area meant a reduced HRR.

• The front of the bookcase, resting on top of that unit, typically fell, changing the location of its

1 3 reams of paper.

• The desk surface bowed progressively as it charred through, and then it collapsed, with separate

sections doing so at differing times. The initial desk collapse probably did not greatly affect its burning rate, but ultimately what was left was a complex rubble pile whose burning would not be

predictable from any knowledge of the original configuration coupled with Cone Calorimeter

data.

• The wall panels collapsed at random times, both inward and outward, typically rather late in the

fire.

Thus, the fiirther one goes out on the HRR curve, past the peak, the less it is predictable by an FDS

calculation that retains the original geometry. Fortunately, the major effects appear to occur well after the desk surfaces collapse and the time when contiguous workstations would become ignited and dominate

the heat release.

J-25
Appendix J

Numerical grids of 10, 20, and 40 cm were used to model the fires and ensure that the model was not

sensitive to grid cell size. Figure J-24 shows a preliminary comparison of the FDS HRR prediction with

the measured values for test 2 (no jet fuel, no inert tiles). The quality of fit is typical of the test series.

12000

10000

8000 -

oi

-(->

6000

4000 -

2000 -

0

Heat Release Rate, Single Workstation

Experiment

FDS

0 500 1000 1500 2000 2500

Time (s)

Figure J-24. Preliminary comparison of predicted and measured HRR values.

For these simulations, the thermal properties of the major materials making up the workstations were

derived from Cone Calorimeter experiments. The carpet and privacy panel were modeled as thermoplastics, that is, the burning rate was assumed to be proportional to the heat flux from the

surrounding gases. The desk was modeled as a charring solid, in which a pyrolysis front propagates

through the material leaving a layer of char behind that insulates the material and reduces the burning

rate. Details of the pyrolysis models can be found in the FDS Technical Reference Guide (McGrattan et

al. 2002). Each feature of the experimental curve was related (using annotations made during the tests and from the video tapes) to specific aspects of the workstation combustion.

There are similarities and discrepancies between the experimental data and this prediction.

• The shape and magnitude of the two curves is encouragingly similar, as is the total heat release (area under the curves).

• The peak HRR occurred sooner in the simulation. In the experiment, the time to peak HRR was

strongly influenced by the melting of the chair plastic onto the carpet. As noted above, this level of detail is not captured in the numerical model.

J-26
Interim Report on Experiments to Support Fire Dynamics

• At long times, the simulation drops to the residual HRR of the 2 MW burner somewhat more

abruptly than does the experimental curve, indicating that it has run out of combustible mass

sooner. The importance of this effect is modest, given the geometric changes during the test (listed above) and the similarity of the total heat released. A more complete analysis will be detailed in the forthcoming documentation report.

J.4 EXPERIMENTS FOR FIRE MODEL VALIDATION

Following the experiments described above and the accordant improvements in FDS, a series of large- scale experiments was conducted in the NIST Large Fire Laboratory between November 4 and December

10, 2003. The six experiments were designed to assess the accuracy with which FDS predicts the fire spread, heat release rate, and thermal environment in a compartment burning multiple workstations in a configuration characteristic of that found in the WTC buildings. In each of these experiments, sets of

three workstations, identical to the generic ones tested in Section J. 3, were burned in a large compartment

(see Fig. J-25). The challenges to the model included varying the location of the ignition burner (and

thus the fire ventilation), adding jet fuel and/or noncombustible material occluding a fraction of the workstations' surfaces, and "rubblizing" the workstations. FDS simulation of each test was carried out before the test was conducted.

The steel-frame experimental enclosure was 10.8 m long x 7.0 m wide x 3.4 m tall (35.5 ft x 23 ft x 11 ft) and was lined with three layers of 13 mm (0.5 in) calcium silicate board (see Fig. J-26). There was a subfloor (not included in the above dimensions) to house instrumentation. The enclosure had openings on

the front mimicking window openings through which fresh air entered and heat and combustion products were emitted. The narrowed openings limited the amount of fresh air that entered the burning enclosure.

Each of three workstations was placed on an isolated platform made of calcium silicate board. The top

surface of each platform was flush with the floor of the compartment. Each platform was supported on

water-cooled load cells, located in the subfloor, to monitor the mass of the workstation throughout the

test. The load cells were the same as those described in Section J.3.

Two of the workstations were contiguous, exemplifying a part of the type of cluster that exists in many

large office spaces. The third workstation was separated from the other two by an aisle, representing a

part of a second cluster. This array was to enable assessment of FDS's ability to replicate two different

modes of cubicle-to-cubicle fire spread: direct flame impingement and radiative ignition from the hot

ceiling layer.

J-27
Interim Report on Experiments to Support Fire Dynamics

Figure J-25. Plan view of test configuration.

J-28
Interim Report on Experiments to Support Fire Dynamics

0.9 m

APPROXIMATE

DIMENSIONS

OPENINGS

2.1 m tall - 0.45 m wide

COLUMNS

0.5 m wide

0.3 m

INSTRUMENTATION SUB-FLOOR (external to experimental enclosure)

FIRE PAN AND

CUBICLE 3 SPRAY BURNER CUBICLE 1 LOCATION 1 Figure J-26. Elevation view of test configuration.

The west end of the enclosure was located under a 10 m x 12m hood for collection of the effluent and

measurement of the heat release rate.

Other instrumentation included:

• Four floor-to-ceiling trees of thermocouples to measure vertical profiles of temperature(see

Fig. J-27);

• Thermocouples on the desk surfaces to track flame spread;

• Two downward-facing, water-cooled Schmidt-Boelter total heat flux gages in the ceiling to measure radiative heat flux;

• Two water-cooled Schmidt-Boelter total heat flux gages mounted on the west wall; and

• Four video cameras placed to record the progression of flame spread over the objects in the

cubicle.

J-29
Interim Report on Experiments to Support Fire Dynamics

J-30
Interim Report on Experiments to Support Fire Dynamics

The liquid spray burner, pan, and fuel (mixture of heptanes) were the same as used in the Section J.

3

experiments. Depending on the test, the burner was located abutting the top of a workstation partition at the east end of cubicle 1 or the west end of cubicle 2. The ignition fire intensity was a nominal 2 MW

fire. The spray burner was operated for the first few minutes of the tests, for either 2 min or 10 min

depending on the test scenario.

Materials

The workstations were of the same type as the generic units used in the experiments reported in Section J.3. The carpet tiles, which covered the floor of the cubicles and the aisle, were also the same

type used in experiments reported in Section J. 3.

Test Variables

The experimental matrix is shown in Table J^. The experiments investigated the impact of several

parameters on the fire behavior:

Table J-4. Test mat rix. Test Ceiling Tiles Jet Fuel Burner Location Workstations Windows

1 None None Front hitact No

2 None None Front Intact No

3 Present Present Front Intact No

4 Present None Rear Intact No

5 Present Present Rear "Rubble" No

6 None Present Rear Intact Yes

• Location of the burner. This was placed either abutting the west end of cubicle 1 or abutting the

east end of cubicle 2. These two sites resulted in significantly different access to the air needed

for combustion. In the former ("fronf) location, much of the oxygen in the air initially entering

the enclosure was consumed by the burner and the burning cubicle 1 , with the resuh that limited

oxygen was available for combustion in the middle and rear of the compartment. With the burner

in the latter ("rear") location, the fresh air passed directly to the rear of the compartment.

• The application of 1 2 L ofjet fuel evenly distributed about each workstation. The procedure was

the same as used in the experiments reported in Section J. 3. The presence of fallen ceiling tiles. Having seen the effect of coverage of 30 percent of the top surfaces in the previous test series,

NIST covered approximately 70 percent of the top surfaces here.

• Fractured furniture. In one experiment (test 5), NIST investigated the effect of different degrees

of "rubblizing" the furniture.

- In cubicle 1 , the workstation pieces were placed unassembled on top of each other,

occupying the same footprint as the assembled workstation. The same mass of

combustibles was present as in the fully assembled cubicle tests. No steel filling cabinets were used. Ceiling tiles and broken up drywall were intermixed with the rubble.

J-31
Appendix J

- Cubicle 2 was the same as cubicle 1 , except without the drywall.

- For cubicle 3, the workstation was partially assembled. The same mass of ceiling tile and

drywall as in cubicle 1 were intermixed with the cubicle components.

- Window breakage. Test 6 had four glass windows mounted on the north end of the west

wall. During the course of the fires in the WTC towers, a number of windows were

broken, presumably by the heat from the fires. These result in a change in both the

degree and pattern of ventilation.

Test Procedure

This was similar to that followed in the experiments reported in Section J. 3, except that the ignition

burner was turned off after approximately 2 min for tests 3, 5, and 6 (when jet fuel was present) and for 10 min for Tests 1, 2, and 4. The tests continued until the HRR fell below 0.5 MW, which was typically

60 min after ignition.

J.4.1 Preliminary Results

Figures J-28 and J-29 show the east view of the compartment before and during a test, respectively. A

few observations about the tests were:

• The peak HRR was approximately 1 1 MW for four of the tests. In test 5, the peak value was only

approximately 6 MW. In test 6, the peak HRR reached almost 16 MW.

• As in the single workstation tests, the peak value was reached earlier when jet fuel was present.

Figure J-30 shows how the enclosure was represented in FDS. The computational grid size was 0.4 m

(1 .3 ft) on a side. Note that the chair, computer monitor, and paper have been collected together into "boxes" with comparable mass to the various items that were spread throughout the workstations. To the

right are five windows that are similar in size to those of WTC 1 and WTC 2.

The preliminary plot in Fig. J-3 1 depicts the degree to which the heat release rate measurements agreed

with those predicted by FDS.

The overall degree of agreement between the model and the experimental data is quite good, despite some

modest local differences. In the analysis leading to candidate improvements in the modeling, it is important to maintain perspective of the effect of these differences on the accuracy needed in reconstructing the actual WTC fires. For fires that are sufficiently severe that they threaten the structural integrity of the building, many such workstations will be burning concurrently. These workstations will

be at various stages of their combustion. Thus, for example, features occurring at long times in Fig. J-3

may not merit closer replication, while those features occurring at short times (and thus have a bearing on

the ease of fire spread among workstations) may merit attention.

J-3 2
Interim Report on Experiments to Support Fire Dynamics

J-33
Appendix J

Figure J-30. Fire Dynamics Simulator 3D rendition of experimental enclosure.

0 500 1000 1500 2000 2500 3000

Time (s)

Figure J-31. Comparison of measured and predicted heat release rate for test 1.

J.5 REFERENCES

McGrattan, K.B., H.R. Baum, R.G. Rehm, G.P. Forney, J.E. Floyd, K. Prasad and S. Hostikka. 2002. Fire

Dynamics Simulator (Version 3), Technical Reference Guide. NISTIR 6783. National Institute of

Standards and Technology, Gaithersburg, MD.

J-34
Attachment 1 Simulating the Fires in the World Trade Center

1.1 INTRODUCTION

In the months following the attacks on the World Trade Center (WTC) and the Pentagon, there was an

active debate in the fire protection engineering community about the fires that erupted following the

impact of the aircraft on the buildings. Because fires of this magnitude in these types of buildings are

rare, there is a wide spectrum of opinion about the fire temperatures and their effect on the structural steel. Much of the fire literature consists of empirical correlations derived from experiments ranging from

bench scale to room scale. Extrapolating these well-known correlations to the WTC requires a reexamination of the underlying assumptions. Many of these correlations are appropriate for a narrow

range of fire sizes and building geometries, and cannot be directly applied to the WTC fire scenarios. As

a result, computer fire models that have been developed over the past decade are being applied to the

analysis.

As part of the investigation, the National Institute of Standards and Technology (NIST) has conducted

simulations of the fires in each building using a computational fluid dynamics (CFD) model known as the

Fire Dynamics Simulator (FDS). This attachment will describe the experiments conducted at NIST to calibrate and validate the FDS model for use in the WTC project, and it will describe the techniques

developed to simulate the very extensive fires that spread over 6 to 12 floors in the different buildings.

1.1.1 Experimental Program

Two large-scale test series were conducted to provide validation for the FDS, plus various small-scale

experiments were conducted to provide the model with input data for different materials. The large-scale

test programs are referred to as Phase 1 and Phase 2. Both test series involved fires in compartments with

the same ceiling height as a floor in WTC 1 or WTC 2. Phase 1 was a series of fire tests with a liquid fuel spray burner generating a fixed amount of energy in a compartment with various targets and obstructions,

like columns, trusses and other steel objects. These tests were designed to test the accuracy of the model,

and its sensitivity to changes in various input parameters. Phase 2 was a series of fire tests in which office

workstations similar to those used in WTC 1, WTC 2, and WTC 7 were burned in a compartment with

limited openings to simulate the under-ventilated conditions of the WTC fires. These tests were designed

to test the model's ability to characterize the burning behavior of real furnishings under conditions typical

of the WTC fires. Only the Phase 2 work will be discussed in this attachment.

1.1.2 NIST Fire Dynamics Simulator

FDS is a CFD model of fire-driven fluid flow. It solves numerically a form of the Navier-Stokes

equations appropriate for low-speed, thermally-driven flow with an emphasis on smoke and heat transport

from fires (McGrattan et al. 2002). Version 1 was publicly released in February 2000. The core algorithm is an explicit predictor-corrector scheme, second order accurate in space and time. Turbulence

is treated by means of the Smagorinsky form of Large Eddy Simulation (LES). For most applications.

J-35
Appendix J

FDS uses a mixture fraction combustion model. The mixture fraction is a conserved scalar quantity that is defined as the fraction of gas at a given point in the flow field that originated as fuel. The model assumes

that combustion is mixing-controlled, and that the reaction of fuel and oxygen is infinitely fast. The mass

fractions of all of the major reactants and products can be derived from the mixture fraction by means of

"state relations," empirical expressions arrived at by a combination of simplified analysis and

measurement.

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; thus the name given to it is the

Finite Volume Method (FVM). Using approximately 100 discrete angles, the finite volume solver

requires about 15 percent of the total CPU time of a calculation, a modest cost given the complexity of

radiation heat transfer. FDS approximates the governing equations on a rectilinear grid. The user

prescribes rectangular obstructions that are forced to conform with the underlying grid.

All solid surfaces are assigned thermal boundary conditions, plus information about the burning behavior

of the material. Usually, material properties are stored in a database and invoked by name by the user. An

extensive effort was undertaken to characterize the thermal properties of common items found in an office

setting, like privacy panels, stacks of paper, computer monitors, office chairs, pressboard tables, desks,

and carpeting. These materials will be described next.

1.2 CALIBRATION AND VALIDATION EXPERIMENTS

The experimental program concentrated on the thermal properties of the office furnishings that constituted the bulk of the combustible fiiel within the WTC buildings under study. Several types of

office workstations typical of those used in WTC 1 and WTC 2 were purchased at area office supply

stores. The thermal properties of the major materials making up the workstations were derived from cone

calorimeter experiments. These properties were input into FDS, which was used to simulate the burning

behavior of a single workstation burning under a 2.5 m ceiling with baffles to contain a hot layer of

smoke above the burning workstation. Other than the baffled ceiling, no walls surrounded the workstation

other than its own privacy panels. The thermal properties of the workstation components were adjusted

slightly so that the FDS prediction of the heat release rate would match the experiment. Then the model was used to predict the heat release rate of 3 workstations burning within a large enclosure. The purpose

of this exercise was to determine if FDS could simulate the dynamics of a fire in a setting similar to WTC 1, WTC 2, and WTC 7.

1.2.1 Description of the Workstation Components

Cone calorimeter experiments at three different heat fluxes were performed for the carpet, desk (wood),

computer monitor, chair, privacy panel, and stacked paper. For the simulations of the WTC fires, only the

carpet, desk and privacy panel data were used directly. The carpet and privacy panel were modeled as thermoplastics, that is, the burning rate is assumed to be proportional to the heat fiux from the

surrounding gases. The desk was modeled as a charring solid, in which a pyrolysis front propagates

through the material leaving a layer of char behind that insulates the material and reduces the burning

rate. Details of the pyrolysis models can be found in the FDS Technical Reference Guide (McGrattan et

al. 2002).

J-36
Interim Report on Experiments to Support Fire Dynamics

The desk was modeled as a charring sohd. The thermal properties of the wood and its char were taken

from both the calorimeter experiments and the work of Ritchie et al. (1997). It is 2.8 cm thick with

density 450 kg/m\ specific heat 1.2 kJ/kg/K at 20 °C and 1.6 kJ/kg/K at 900 °C, conductivity

0.13 W/m/K at 20 °C and 0. 16 W/m/K at 900 °C. The ignition temperature is 360 °C and the heat of

combustion is 14,000 kJ/kg ± 800 kJ/kg. Its total available energy content is 210 MJ/m^ ± 50 MJ/m".

The carpet was modeled as a thermoplastic with density 750 kg/m^, specific heat 4.5 kJ/kg/K,

conductivity 0.16 W/m/K, ignition temperature 290 °C, thickness 6 mm, heat of vaporization 2,000 kJ/kg,

heat of combustion 22,300 kJ/kg ± 600 kJ/kg, and total available energy content 61 MJ/m' ± 2 MJ/m".

The privacy panel was modeled as a thermally-thin thermoplastic. The product of specific heat, thickness

and density is 0.73 kJ/m"/K. Its surface density is 0.25 kg/m", ignition temperature 380 °C, heat of

vaporization 6,000 kJ/kg, heat of combustion 30,000 kJ/kg ± 500 kJ/kg. Its total available energy content

is 6.0 MJ/m- ± 1.3 MJ/ m".

The test compartment walls and ceiling were made of three layers of 1.27 cm (0.5 in) thick Marinite I, a product of BNZ Materials, Inc. (http://www.bnzmaterials.com).^ The manufacturer provided the thermal

properties of the material used in the calculation. The density is 737 kg/m^ conductivity 0.12 W/m/K.

The specific heat ranged from 1 .2 kJ/kg/K at 93 °C to 1 .4 kJ/kg/K at 425 °C.

In the simulations of the fires within the WTC, the chair, computer, paper, and other miscellaneous items

within the workstation were modeled as a single item by lumping the mass together into large "boxes"

and distributing them throughout the workstation. It is common practice in fire protection engineering to use surrogate materials for fire experiments, and this practice has been extended to numerical modeling.

Over the years, various items have been developed that are representative of various types of

commodities. For example, wood cribs are often used to represent ordinary combustibles found in residential or light industrial settings. Paper cartons with various amounts of plastic within are also used

as surrogates for a wide range of retail commodities. One in particular is called the FMRC (Factory

Mutual Research Corp.) Standard Plastic Commodity, or more commonly. Group A Plastic. This test fuel

is often used in sprinkler approval testing at Factory Mutual and Underwriters Laboratories in the US, and

similar test fuels have been developed in Europe. In the late 1 990s, FDS was used to simulate large scale rack storage fires to determine the effectiveness of the combined use of sprinklers, roof vents and draft curtains (curtain boards). As part of this effort, a considerable amount of work was done to characterize

the thermal properties of Group A Plastic (Hamins and McGrattan 2003). Because Group A Plastic has

been shown to be fairly representative of fires fueled by a mixture of paper (cellulosic materials) and

plastic, and because it has been used in numerous FDS simulations, it was decided to model the contents of the office workstations with a fuel similar to Group A Plastic. Blind predictions of the single open

workstation bums were made using the material properties obtained during the sprinkler/roof vent study,

and then these properties were adjusted to match the results of the experiments. Thus, the single

workstation bums served to calibrate the model. They were not intended to be validation experiments.

The validation experiments consisted of buming 3 workstations at a time in an under-ventilated

compartment.

^ Certain commercial equipment, instruments, or materials are identified in this document. Such identification does not imply

recommendation or endorsement by the National Institute of Standards and Technology, nor does it imply that the products

identified are necessarily the best available for the purpose.

J-37
Appendix J

The surrogate fuel is modeled as a homogenous solid whose density is 172 kg/m^ The paper carton is treated as a thermally-thin material whose density x specific heat x thickness is 1 .0 kJ/m"/K. Its ignition

temperature is 370 °C and the heat of combustion is 30,000 kJ/kg. The heat release rate of the boxes

ramps up to 450 kW/m' in about 1 min. Note that this fuel package is similar, but not the same, as Group

A Plastic. The density has been increased to account for all the miscellaneous items within the

workstation. Also note that unlike the desk, partition and carpet, the boxes are simply given a burning

rate rather than a heat of vaporization, meaning that the boxes will bum at the given rate regardless of the

heat flux upon them as long as the surface temperature remains above its ignition temperature. The

reason for this is that it is not possible to predict the burning rate using the heat feedback approach

because the geometry of the scattered fuel is too complex to directly predict the response of the material

to the thermal insult. By collecting all the scattered items into boxes, the geometry of the combustibles is greatly simplified, and as a result the burning behavior must be simplified as well.

1.2.2 Description of the Simulations

The geometry of the compartment is relatively simple. The overall enclosure is rectangular, as are the vents and most of the obstructions. Numerical grids of 20 and 40 cm were used to model the fires. The

purpose of the grid variation was to ensure that the model was not sensitive to the change in grid cell size. Typically, enclosures of this size are modeled using 10 cm grid cells. However, for the simulations of

WTC 1, WTC 2, and WTC 7, a 40 cm grid was used. By simulating the experiments at 20 cm and 40 cm,

NIST can test if the model produces significantly different results with grid cells of different sizes. Figure 1-1 is a snapshot of a simulation showing the fire and the major geometric features of the compartment for the simulations. Note that the surrogate fuel packages are placed roughly where the computer monitor, chair and paper were located. Six tests were performed, with various ignitor locations

and fuel arrangements. A 2 MW burner was pi aced either near the windows of the compartment

overlooking the workstation nearest the openings in Tests 1, 2, and 3. The burner was placed towards the rear of the compartment overlooking the workstation in the rear of the compartment in Tests 4, 5, and 6. In Tests 3,5, and 6 Jet A fuel was poured over the workstations and surrounding carpet. To simulate this in the model, spray nozzles were positioned over the center of each workstation, 2 m above the floor. These nozzles are normally used to simulate water sprinklers, but in this case, the water was replaced by a

liquid having similar properties to Jet A. The nozzles were activated for 2 s, in which time the equivalent

amount of liquid as in the tests was ejected and spread over the furnishings.

In Test 5, Workstations 1 and 2 were disassembled prior to the bum and the contents were piled on top of

the respective load cells. To model this scenario, the burning rate of the collective fuel packages was

reduced by one half to account for the decrease in buming area of the fuel pile. The choice of one half was somewhat ad hoc. No free bums of workstation parts had been performed. This was the only test in which the simulated fuel packages had to be modified txom their free-bum values. In this regard, Test 5 was used to calibrate, not validate, the model.

J-38
Interim Report on Experiments to Support Fire Dynamics

NET

Figure 1-1. Geometry of the Phase 2 simulations.

From a modeling perspective, the objective of the simulations of the Phase 2 experiments was to demonstrate that a simplified model of an office workstation can be used to predict the burning behavior

of a group of workstations in an enclosure with features similar to WTC 1, WTC 2, and WTC 7. Because

of the magnitude of the simulations of the building fires, the model of the workstation had to be fairly

crude. However, because of the many uncertainties in the initial conditions of the fire simulations, the

lack of detail in the model is not considered to be a problem. The model fires had similar growth patterns,

peak heat release rates, decay patterns, and compartment temperatures.

The model also captured the major features of the individual tests. For example. Tests 1 and 4 were

similar in design except for the burner location. In Test 1 the burner was near the windows; in Test 4 it was near the rear of the compartment. The peak heat release rate was reached in about 15 min in Test 1, whereas it was reached in about 10 min in Test 4. The model shows a similar trend. The faster growth of

Test 4 is probably due to the fact that the compartment heated up more quickly with the fire deep inside

rather than near the windows, leading to more rapid spread of the fire across the pre-heated furnishings.

Even though ceiling tiles were distributed over the desk and carpet in Test 4, this did not seem to have a noticeable effect on the growth, or at the very least the burner position seemed to have a far greater role in explaining the difference between Tests 1 and 4. The comparison of HRR between model and experiment

is shown for Test 1 in Figs. 1-2. The upper layer temperature in the rear of the compartment for this same test is shown in Figs. 1-3. The results for the other tests are comparable. The peak HRR and

temperature are predicted well, as well as the duration of the fires. Both the peak values and the duration

of the burning are important for the WTC simulations because it is not only important to predict the temperatures that the structural steel was exposed to, but also the duration of the exposure.

J-39
Appendix J

ISa;

t/3

16000

14000 -

12000 -

10000 -

8000 -

6000 -

4000 -

2000 -

WTC Phase 2, Test 1 Heat Release Rate

Experiment

FDS

500 1000 1500 2000

Time (s)

Figure 1-2. Comparison of HRR for Phase 2, Test 1

2500 3000

1200

B

CD H

Figure 1-3. Upper layer gas temperature in the rear of the compartment,

Phase 2, Test 1.

J^O
Interim Report on Experiments to Support Fire Dynamics

1.3 SIMULATIONS OF THE FIRES IN WTC 1 AND WTC 2

This section describes how the physical geometry of the buildings was described in the numerical model.

Information about the layout of the relevant floors was obtained from architectural drawings provided by

the occupants. For floors where information was not available, the geometry of a nearby floor or a floor of

similar use was substituted. Information abqut exterior damage and window breakage was obtained by

studying thousands of photographs and videos. There was no attempt made to predict the window

breakage in the simulations. This information was provided as a boundary condition.

1.3.1 Numerical Grid

The windows in WTC 1 and WTC 2 were nominally spaced 1 m apart. In addition, the external columns

plus their aluminum cladding were assumed to be 0.5 m wide. The slab-to-slab floor spacing was

assumed to be 3.6 m. Because of these approximations, a uniform numerical mesh consisting of cells whose dimensions were 0.5 m x 0.5 m x 0.4 m was used. In the model, each tower face consisted of

58 windows, 61 columns, and two 0.5 m spacers next to each comer column. In the real tower geometry,

these spacers formed the bevel. Figure 1^ shows a single floor of the WTC 1 as it is approximated by

the numerical model.

The numerical grid for each floor of WTC 1 and WTC 2 was of dimension 128 x 128 x 9 cells. The

128 cells in the horizontal directions allow for several meters of simulation outside of the external walls. The calculations were run in parallel, thus each floor was assigned to a different processor. The floor

slabs, core walls, and workstations were approximated as thin obstructions. As described in the previous

section, the contents of each workstation were collected into boxes and distributed throughout.

Penetrations in the floor slabs representing elevator shafts and HVAC ducts were created in the model by

defining rectangular plates on top of the floor slab that were removed at the start of the calculation. This

served to carve out holes in the floor. Window breakage was modeled by removing thin obstructions

serving as windows at times obtained from the analysis of photographs and videos. Broken external

columns were removed the same way.

1.3.2 Parallel Processing

Modeling the fires on multiple floors of WTC 1 and WTC 2 is computationally intensive, both in terms of

CPU time and memory. Up to this point in its development, FDS has been limited to calculations small

enough to run on a single CPU and fit into the memory of a desktop personal computer. The WTC study

is an example of a large-scale fire modeling problem that is impossible to analyze without the use of

parallel processing. In terms of parallelization, the exact details of FDS are not important. The approach

taken to run the code on a cluster of machines can be applied to virtually any CFD code, in particular

those that involve three spatial dimensions and time. In such cases, the computational demand is fairly

well represented by the product of the number of computational grid cells and the number of time steps

taken to advance the solution of the governing equations in time. For example, if the computational grid

consists of 1 million cells and the simulation requires ten thousand time steps, the demand is

lO'" cell-cycles. The overall demand can be broken down into memory requirement and CPU time. The

memory requirement is a function of the number of grid cells; the CPU time is a function of the number

of time steps.

J-41
Appendix J

Figure 1-4. Plan view of a typical floor in WTC 1.

Roughly speaking, state of the art 32 bit processors can complete roughly 100,000 FDS cell-cycles per

second. Realistic simulations of fires such as those in the WTC require on the order of 500,000,000,000

cell-cycles, or about two months of calculation on a single 2 GHz processor. Plus, the calculation would

require 6 to 12 gigabytes of memory (RAM), well over the 4 gigabyte address space of 32 bit processors. Because of this, the WTC calculations are not only impractical on single processor systems, they are impossible on any 32 bit processor. A 64 bit processor system may theoretically handle the static memory

requirements of a large simulation time, but the run times for large calculations remain prohibitive.

Because of the computational and memory issues of large fire simulations, a parallel version of the fire model must meet the two fundamental requirements discussed above, as well as satisfy a number of

practical implementation concerns. Both the computational and the memory requirements must be

distributed across muhiple processors. The simulation must be done so that each processor uses less than

J^2
Interim Report on Experiments to Support Fire Dynamics

4 gigabytes of memory, while enough processors must be used to reduce the simulation to a practical

length of time, of the order of one week.

Because the computational load is distributed throughout much of the source code, NIST has chosen to break up the calculation into multiple spatial blocks, with each block essentially doing the same type of

calculation. A feature common to most CFD codes is multi-block or multi-mesh structure in which more

than one structured grid is used in the calculation. This feature is exploited by simply putting the data and

computation for each block on a different processor. This has advantages and some limitations. The

advantages are ( 1) a natural and scalable extension of the existing code, (2) the amount of data

communication will be kept to a minimum, since only overlap information needs to be communicated,

rather than the data for full blocks, (3) source code changes are localized in small communication

routines, (4) development is fairly fast. The disadvantages to the multi-block approach are (1) equal

distribution of work across processors (load balancing) depends on spatial symmetry in the simulation,

such as the translationally synmietric geometry of the WTC floors, (2) the level of parallelism and the

speed up of the calculation is limited to the number of spatial blocks that can be used in the calculation.

These limitations are not severe in many cases, including the WTC.

Because NIST is interested in a scalable, portable code. Message Passing Interface (MPI) is used. This is a standard, well-documented system of implementing parallel processing, that can work with shared

memory, distributed memory, or combinations of those architectures (Gropp, 1999). Our goal in using

MPI was to produce a code that, except for the requirement of the MPI library, would be as portable and

standardized as the sequential version. The parallel code runs on most computer platforms, including

networked Windows-based PCs. NIST opted for a cluster of commodity personal computers running

Linux, cormected by a gigabit ethemet network. The individual processors are in the range of 2.0 to 2.8 GHz, and dual processor machines were chosen to save space and to allow the addition of OpenMP

code as a future extension to the MPI-based code. For production work in the NIST laboratory, two

clusters are used: a smaller, development cluster to develop and debug the code, and a larger cluster with

128 processors. Using both clusters provides the capability to run six to eight large parallel processing

jobs simultaneously.

1.3.3 Sample Simulation

Shown in Fig. 1-5 is a sequence of snapshots showing the predicted upper layer temperatures on a floor of WTC 1 at time increments of 15 min. The first image is a cut-away showing the damage to the north

face of the tower and the layout of the walls and furnishings. The subsequent images are color coded by

temperamre, with the red (or dark) patches representing temperatures in the vicinity of 1,000 °C. Initially,

these hot areas of active burning are near the impact zone at the north end (foreground of picture), but

migrate towards the south as the combustible furnishings are exhausted. Driving the progress of the fires

is the breaking of windows that provide air to the oxygen starved fire. The window breakage is not

predicted by the model; it is an imposed boundary condition resulting from the analysis of thousands of

photographs and videos recorded that day by eye witnesses. The uncertainty in the window break times is on the order of 5 min in areas not obscured by smoke.

J-43
Appendix J

Figure 1-5. Predicted upper layer temperatures of a floor of WTC 1.

The burning behavior shown by the simulation is similar to that of the fires in experiments conducted by

Ian Thomas and Ian Bennetts (1999). They looked at fire spread in long and wide enclosures with a

single ventilation opening, where the fires were ignited at various points deep within the bench-scale

compartments used. The fires would rapidly spread across the liquid or solid fuels covering the floor without consuming much of the fuel. The fires would then surround the compartment opening and bum

back into the compartment as the fuel near the opening was exhausted. In the WTC simulations, fires are ignited over a wide area by simulated spray nozzles ejecting a liquid with properties of aircraft fuel. Much of the available oxygen is consumed rapidly, driving the fires to the openings created by the

aircraft. The fires move away from the initial impact area as the nearby furnishings are exhausted, and as windows are broken out away in other parts of the building.

J^4
Interim Report on Experiments to Support Fire Dynamics

1.4 SUMMARY

The investigation into the cause of the collapse of WTC 1, WTC 2, and WTC 7 by NIST will not be

completed until the fall of 2004. Work is on-going to simulate the weakening of the structural steel due to the aircraft impacts and the fires. Nevertheless, the fire experiments and simulations performed to date

have improved our ability to analyze the response of any large building or structure to fire. In the years

ahead, these techniques will become increasingly widespread due to faster computers and the ability to harness an entire set of off-the-shelf personal computers to perform very large calculations. Effective

modeling is a combination of fast computers, efficient algorithms, and well-planned small and large scale experiments to provide both input to the model and a validation of results. Projects as complicated as the

WTC study are rarely conducted using modeling alone. There is and will always be a need to coordinate

computation and experiment to reconstruct the dynamics of large fires.

1.5 REFERENCES

Gropp, W., E. Lusk, and A. Skjellum. 1999. Using MPI, MIT Press, Cambridge, Massachusetts, USA.

Hamins, A., and K.B. McGrattan. 2003. Reduced-Scale Experiments on the Water Suppression of a RackStorage Commodity Fire for Calibration of a CFD Fire Model. Fire Safety Science: Proceedings of

the Seventh International Symposium. International Association for Fire Safety Science.

McGrattan, K.B., H.R. Baum. R.G. Rehm, G.P. Forney, J.E. Floyd, K. Prasad and S. Hostikka. 2002. Fire

Dynamics Simulator (Version 3), Technical Reference Guide. NISTIR 6783. National Institute of

Standards and Technology. Gaithersburg, MD.

Ritchie, S.J., K.D. Steckler, A. Hamins, T.G. Cleary, J.C. Yang, and T. Kashiwagi. 1997. The Effect of

Sample Size on the Heat Release Rate of Charring Materials. Fire Safety Science: Proceedings of the

Fifth International Symposium. International Association for Fire Safety Science.

Thomas, I.R. and I.D. Bennetts. 1999. Fires in Enclosures with Single Vendlation Openings-Comparison

of Long and Wide Enclosures. Fire Safety Science: Proceedings of the Sixth International

Symposium. International Association for Fire Safety Science.

J-45
Appendix J

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J^6
Table of Contents

List of Figures K-iii

List of Tables K-vi

Appendix K

Interim Report on Subsystem Structural Analysis of the WTC Towers K-1

K.l Purpose K-1

K.2 Scope of Work K-1

K.3 Description of Subsystem Structures K-2

K.3 . 1 Full Floor Subsystem Description K-2

K.3. 2 Exterior Wall Subsystem Description K-3

K.4 Loads K-^

K.5 Materials K-4

K.5.1 Concrete K-5

K.5.2 Steel K-8

K.5.3 Welds K-15

K.5.4 Bolts K-15

K.5. 5 Coefficient of Friction K-16

K.5.6 Symbols K-16

K.6 Model Conversion From SAP to ANSYS K-1

7

K.6.1 Translation Procedure K-18

K.6.2 Challenges K-19

K.6.3 Status K-19

K.7 Full Floor Subsystems K 24

K.7. 1 Full Floor Model K-24

K.7.2 Knuckle Analysis K-26

K.7. 3 Column Truss Seats K-3

1

K.7.4 Modeling Connection Failure by Break Elements K-40

K.7. 5 Truss Model K--^2

K.8 Exterior Wall Subsystem K-57

K.8.1 Description of Exterior Wall Subsystem Model K-61

K.8.2 Validation of the Exterior Wall Subsystem Model K-61

K.8.3 Model of One-Story High Exterior Column K-6

1

K-i
Appendix K

K.8.4 Model of Nine-Story High Exterior Model K-61

K.8.5 Models of the Column Splice K-61

K.8.6 Prefabricated Panel Model K-61

K.8.7 Ongoing Work on the Exterior Wall Subsystem Model K-61

K.9 Floor Truss Dynamic Response Due to Impact of Dropping Floor K-61

K.9.1 Impact of Dropping Floor K-61

K.9. 2 Purpose and Scope K-61

K.9. 3 Method of Analysis K-61

K.9.4 Results K-61

K.9. 5 Conclusions K-61

K.IO References K-61

K-ii
List of Figures

Figure K-1 . Temperature-dependent concrete properties K-6

Figure K-2. Concrete stress-strain curves K-7

Figure K-3. Temperature-dependent properties for all steels K-9

Figure K^. Stress-strain relationships for Material ID 1 steel K-1

1

Figure K-5. Creep behavior at elevated temperatures for Material ID 1 steel K-1

3

Figure K-6. Maximum plastic strain from the finite element analysis and limiting

plastic strain K 14

Figure K-7. 7/8 in. A325 bolt load-elongation curves at elevated temperatures K-1

5

Figure K-8. Converted ANSYS model for floor 96 of WTC 1 : overall view K-20

Figure K-9. Converted ANSYS model for floor 96 of WTC 1 : partial view near comer

of building K-20

Figure K-10. Converted ANSYS model for floor 96 of WTC 1 : close-up view at comer

of building K-21

Figure K-1 1 . Converted ANSYS model for floor 96 of WTC 1 : view of floor beams

and columns K-21

Figure K-1 2. Deformed shape of gravity load case for SAP floor model K-22

Figure K-1 3. Deformed shape of gravity load case for ANSYS floor model with BEAM44

(Euler beam) elements K-22

Figure K-1 4. Dominant mode shape of floor stmcture for SAP floor model K-23

Figure K-1 5. Dominant mode shape of floor stmcture sase for ANSYS floor model K-23

Figure K-1 6. Transverse shear test of a knuckle K-27

Figure K-1 7. Longitudinal shear test of a knuckle K-2

7

Figure K-1 8. Finite element models of knuckle shear tests K-29

Figure K-1 9. Compressive stresses in longitudinal shear finite element model K-29

Figure K-20. Compressive stresses in transverse shear finite element model K-30

Figure K-2 1 . Shear force versus displacement from finite element model for longitudinal

shear of two knuckles K-30

Figure K-22. Shear force versus displacement from finite element model for transverse shear of two knuckles K 30

Figure K-23. Interior seat K-32

Figure K-24. Exterior seat K-32

Figure K-25. Tmss seat detail location on northeast quadrant of floor 96 of WTC 1 K-33

Figure K-26. Finite element model of exterior seat K-35

Figure K-27. Failure sequence of the exterior seats against tensile force K-36

K-iii
Appendix K

Figure K-28. Typical tensile force resistance from exterior seat K-36

Figure K-29. Strength of combined vertical and horizontal force K-38

Figure K-30. Basic mathematical model of connection failure K—41

Figure K-3 1 . Simplified model of interior seat K-43

Figure K-32. Results of simplified seat model capturing failure from truss walking off

interior seat K-44

Figure K-33. Results of simplified seat model capturing failure from exceeding the interior seat vertical shear capacity K-A4

Figure K-34. Simplified model of exterior seat K-45

Figure K-3 5. Simplified model of knuckle

Figure K-36. Truss model K-48

Figure K-37. Boundary conditions K-50

Figure K-38. Break elements at the interior end of slab K-50

Figure K-39. Comparison of resistance weld strength and yield strength of web member at elevated temperatures K-52

Figure K-40. Displacement versus temperature K-53

Figure K^l. Axial force in truss members versus temperature K-54

Figure K-42. Axial stress contour in the truss members at 663 "^C K-55

Figure K-43. Force in the knuckles versus temperature K-55

Figure K^4. Reaction forces at seats K-56

Figure K-45. Finite element analysis results from increasing gravity K-56

Figure K^6. Exterior wall subsystem structure K-58

Figure K-47. Column and floor number materials and splice types K-59

Figure K-48. Spandrels and spandrel splices K-60

Figure K^9. Exterior wall subsystem model, viewed from inside of WTC 1 K-61

Figure K-50. Portion of exterior wall subsystem model showing number of elements used K-61

Figure K-5 1 . Schematic representation of columns used in the exterior wall

subsystem model K-61

Figure K-52. One-story exterior column model K-61

Figure K-53. Load-deflection of column at room temperature and 700 °C K-61

Figure K-54. Local buckling of column at room temperature K-61

Figure K-55. Plastic hinge in column at room temperature K-61

Figure K-56. Deformed shape of column at maximum axial load at 700 °C K-61

Figure K-57. Nine-story column model K-61

Figure K-58. Variation of maximum temperature and corresponding yield stress with time,

fire scenario G K-61

K-iv
Appendix K

Figure K-59. Maximum compressive and tensile axial stress and corresponding yield stress with time, fire scenario G K-61

Figure K-60. Deformed shape of column at 400 s, 3,200 s, and 5,000 s (floors 95 97) K-61

Figure K-61. Plate model of column splice, floors 97-98 K-61

Figure K-62. Simplified model of column splice K 61

Figure K-63. Column splice details, plate model and simplified model K-61

Figure K-64. Variation of axial displacement with axial load K-61

Figure K-65. Variation of lateral displacement with shear load K-61

Figure K-66. Variation of rotation with moment, out of wall plane K-61

Figure K-67. Variation of rotation with moment, in plane of wall K 61

Figure K-68. Variation of twist angle with torque K-61

Figure K-69. SAP2000 model of prefabricated panel K-61

Figure K-70. ANSYS model of prefabricated panel showing geometry and number of

elements used K-61

Figure K-7 1 . ANSYS model of prefabricated panel showing loading and

boundary conditions K-61

Figure K 72. Deflection of prefabricated panels under 100 kip lateral load K-61

Figure K-73. Deflection of prefabricated panels under 100 kip transverse load K-61

Figure K-74. Deflection of prefabricated panels under 10 kip vertical load K-61

Figure K-75. Schematic of full truss or partial truss drop and diagonal crushing at impact K 61

Figure K-76. Target truss resistance against increasing acceleration

K-v
List of Tables

Table K-1 . Steel types used in WTC 1 and WTC 2 K-9

Table K-2. Parameters for k(T) and n(T) K-1

Table K-3. Uniaxial plastic strain at fracture K-1

Table K-4. Comparison of SAP and ANSYS results for gravity load case K-1

9

Table K-5. Comparison of SAP and ANSYS Modal Analysis Results K-23

Table K-6. Material properties used for truss seat calculations K-34

Table K-7. Interior seat capacity against vertical force K-37

Table K-8. Interior seat capacity against tensile force K-3

8

Table K-9. Exterior seat capacity against vertical force K-39

Table K-10. Exterior seat capacity against horizontal tensile force K-39

Table K-1 1 . Compression strength of gusset plate K-40

Table K-1 2. Material assignments in truss model K-51

Table K-1 3. Resistance weld strength K-5

Table K-14. Column sectional properties K-60

Table K-1 5. Spandrel splice details K-60

Table K-1 6. Column splice details K-60

Table K-1 7. Validation results K-6

Table K-1 8. Demand-to-Capacity ratio of long-span truss for static gravity load K-61

Table K-1 9. Demand-to-Capacity ratio of long-span truss for dynamic impact load from

full truss drop K-61

Table K-20. Peak deflection response due to static gravity and dynamic impact K-61

K-vi
Appendix K

Interim Report on Subsystem Structural Analysis of the

WTC Towers

K.1 PURPOSE

Project 6 addresses the first primary objective of the technical investigation led by the National kistitute of Standards and Technology (NIST) of the World Trade Center (WTC) disaster: to determine why and

how the WTC towers (WTC 1 and WTC 2) collapsed following the initial impacts of the aircraft. Specifically, the objectives of this project are to determine the response of the structural components and

systems to the fire environment in WTC 1 and WTC 2 and to identify probable structural collapse

mechanisms. This appendix documents the progress achieved to date on Project 6 in thermal/structural

modeling of WTC 1 and WTC 2.

Project 6 seeks to determine the response of structural components and systems to the fire environment in the WTC 1 and WTC 2 and to identify probable structural collapse mechanisms by ( I ) evaluating the response of floor and column systems under fire conditions, (2) evaluating the response of the WTC

towers without and with aircraft impact damage under fire conditions, (3) conducting tests of structural components and systems under fire conditions, and (4) evaluating competing failure hypotheses for the

WTC towers.

K.2 scope of work

The scope of the work consists of the following three tasks:

• Task 1, Subsystem Structural Analysis. The objective of Task 1 includes structural analysis of components and two subsystems, a full-floor subsystem, and an exterior wall

subsystem. Task 1 is intended to provide guidance for the development of the global finite element models (FEMs) with respect to element types and sizes, appropriate constitutive

models, and failure criteria for any given structural component. The subsystem analyses also

will help to validate the accuracy of the global analyses, and correlate the results of the fine mesh component analyses with the coarser mesh global analyses of Task 2 and Task 3.

• Task 2, Global Analysis of the WTC Towers' Response to Fire without Impact Damage.

The objectives of this task are to determine the general vulnerability of the towers to

fire-initiated collapse and the role of fire in the towers with respect to structural stability,

sequential failures of components and subsystems, and collapse initiation for the towers

without impact damage.

• Task 3, Global Analysis of the WTC Towers' Response to Fire with Impact Damage.

The objectives of this Task are to determine the relative roles of the impact damage and fires

in the towers with respect to structural stability and sequential failures of the components and

subsystems and to determine probable structural collapse initiafion sequences.

K-1
Appendix K

Work under Task 1 includes the following:

• Develop and validate ANSYS models of the full floor and exterior wall subsystems.

• Evaluate structural responses for the following loading conditions. - Service loads due to gravity (dead and live loads).

- Elevated structural temperatures.

• Identity the possible, likely and most likely failure modes and failure sequences, and the associated temperatures at failure and times-to-failure.

• Identify the changes in mechanical properties or geometry at initiation of component and

subsystems collapse.

• Identify simplifications for the global structural models and/or analyses of subsystem models

to use in Task 2 and Task 3.

The scope of this report is to present the progress made in Task 1 work.

K.3 DESCRIPTION OF SUBSYSTEM STRUCTURES

The full floor subsystem modeled is floor 96 of WTC 1 . The model is believed to be typical in the upper

floors of both towers. The exterior wall subsystem is a nine-column (three-panel) wide by nine-story

(three-panel) high section of the WTC 1 between floor 91 and floor 100 and column 150 and column 158. This area is typical of the exterior walls of the towers and connects to a part of the floor system near the comer with different types of trusses.

K.3.1 Full Floor Subsystem Description

Floor 96 of WTC 1 was identified as an office floor with typical floor construction and loading and,

therefore, was selected as the basis for this model. Components of the floor subsystem are examined for performance under loads and elevated temperatures in different possible failure modes. Understanding of

these component behaviors is used to define the floor models for global analyses of WTC 1 and WTC 2.

The full floor subsystem of floor 96 of WTC 1 includes both office area and core area horizontal framing,

as well as columns immediately above and below this floor.

The floor support in the office area consisted of pairs of steel floor trusses (nominally 60 fit in north-south

and 36 ft in east-west directions) that span between exterior walls and the central core at 6 ft 8 in. on

center. Each of these primary trusses consisted of top and bottom chords fabricated from steel angles and

diagonals fabricated from round bars that extended 3 in. above the top chord at the panel points into the concrete slab in the form of a knuckle. The top chords of the primary trusses were supported at the

central core by truss seats connected to a steel channel that ran continuously between the core columns.

Each pair of trusses was connected to this channel with a seat that included two 1 3/4 in. long slotted holes and two 5/8 in. bolts (one bolt in each truss) as shown in Fig. K 25. Note that the floor truss was

not welded to the seat support.

K-2
Interim Report on Subsystem Structural Analysis of the WTC Towers

At the exterior wall, the truss pair was supported by a seat angle and fastened with two 5/8 in. diameter

bolts in 2 in. long slotted holes. In addition, a gusset plate welded to the spandrel and to the truss top

chord tied the supporting column to the truss, and a pair of straps welded to the top chord and to adjacent

columns tied those columns into the primary trusses. Primary trusses were interconnected by a transverse bridging system consisting of bridging trusses and bridging angles. These bridging trusses were of

similar construction to the primary trusses, although the knuckles for the diagonals did not project above

the top chords. The top chord of the bridging trusses sat 1 1/2 in. below the top chord of the primary

trusses and provided support for the 1 1/2 in., 22 gauge steel deck and the 4 in. thick lightweight concrete

slab. At each comer of the building core, a 36 ft long transfer truss extended out from the comer core column to the exterior wall and supported the 60 ft long primary tmsses. The core area floor consisted of

a 5 in. thick normal-weight concrete slab on 1 1/2 in., 22 gauge steel deck, supported by wide flange

girders and beams connected to the core columns.

Task 1 analyses use the nominal dimensions and design details shown on the drawings, without

modifications resulting fi-om any constmction deviations or tenant modifications. Those modifications

are considered in subsequent Task 3 analyses, which are based on the reference model developed by

Leslie E. Robertson Associates (LERA) under a contract for Project 2. Material properties are based on

information provided by Project 3.

K.3.2 Exterior Wall Subsystem Description

Each side of the towers' exterior wall consisted of fifty-nine 14 in. square box columns spaced at 3 ft 4 in. on center, with 52 in. deep spandrel plates at each floor level. The exterior wall was constmcted from

shop-welded prefabricated panels, each consisting, in general, of three columns and three spandrel beams,

13 ft 4 in. wide by 36 ft high. Except at mechanical floors, the base and top of the stmcture. vertical splices in prefabricated panels were staggered such that within any story, every third prefabricated panel

had a vertical splice. Exterior column splices at the upper stories typically consisted of four 7/8 in. diameter ASTM A325 bolts fastened through the welded butt plates at the tops and bottoms of adjoining

columns. Special prefabricated panels existed for the mechanical floors where no stagger existed at floors

7, 41, 75, and 108. At these mechanical floors, the column splice detail included supplemental field welding in addition to the bolted connection. Horizontal (spandrel-to-spandrel) connections between

prefabricated panels were all field-bolted using splice plates. Comer panels that connected the orthogonal

walls at comers were two-stories tall (24 ft) and consisted of two columns, two spandrel plates, and a third column midway between the two columns on altemate floors.

Various grades of steel, having yield strengths ranging between 42 ksi and 100 ksi. were specified to fabricate the perimeter column and spandrel plates. However, fewer grades were actually used with

somewhat coarser gradation in yield strength than specified. Plate thicknesses also varied, both vertically

and around the building perimeter. Plate thicknesses in the exterior wall were as thin as 1/4 in. at the upper stories, and increased toward the base of the building. The specified plate thicknesses and material

yield strengths differed between the two towers, among NS and EW directions and through the height of

the tower.

An exterior wall subsystem model, nine columns wide and nine floors high, was selected to study the stmctural behavior and failure modes of the exterior wall system. This subsystem model represents the

exterior wall of WTC 1 between floors 91 and 100 and includes column lines 150 through 158. This area

K-3
Appendix K

is located near the comer of the tower (column 159 is at the comer of the north face of WTC 1 . See

Fig. K-27.) The wall subsystem allows evaluations of the interaction of the wall subsystem with thermal

expansion of the floor near the comers. It also connects to various types of tmsses with different

behaviors.

K.4 LOADS

The subsystems and components are analyzed for Dead (D), Live (L), and thermal (Ta) loads. The dead

load consists of structural weights and superimposed dead loads. The superimposed dead loads for floors outside the core consist of the weights of ceiling, mechanical and electrical, fireproofmg, and floor finish,

estimated at 8 psf The superimposed dead load and design live load are defined in the World Trade

Center Design Criteria (LERA 2001). Twenty five (25) percent of the design live load is selected as a reasonable approximation of the load that likely existed at the time of the collapse. (For example,

25 percent of the design live load results in a load of 13.75 psf for the long-span tmsses in the two way

zone of floor 96 with 55 psf design live load.) The service dead and live loads are applied first, followed

by the thermal loads.

The dead and live loads are defined as weights, so that during the collapse process, the gravity loads

remain acting on the stmcture. The weight of debris from the plane will be considered where provided by

Project 2.

The thennal loads, Ta, are temperature time histories for all stmctural members provided by Project 5 for the standard test fire ASTM E 1 1 9 and between three and five representative building fire scenarios of

different intensities and three fire protection conditions.

For analysis of some of the components, discrete values of temperature or temperature distributions in the

form of a ramp from 0 °C to 700 °C (or to a temperature below 700 °C that results in the failure of the component) over 0.5 h followed by a constant temperature of 700 °C for another 0.5 h are used. Failure

modes of the components are evaluated at room temperature and at different elevated temperatures, as

failure modes and failure loads may change with increasing temperature.

Although wind may have had a minor role in the collapse of the towers, Task 1 analyses do not include

wind load effects.

K.5 MATERIALS

The mechanical properties of both steel and concrete are affected significantly by temperature. In the

following sections, the material properties used in this project are specified as a function of temperature.

A material properties catalog is prepared and made accessible to all analysis models. For use in ANSYS,

each material is identified with a number; steels are Material ID 1 through Material ID 29, and concretes

are Material ID 51 through Material ID 83.

K-4
Interim Report on Subsystem Structural Analysis of the WTC Towers

K.5.1 Concrete

Aggregate Types

Two types of concrete were generally used for the flooring inside the towers; lightweight concrete was

used in the office areas, and normal-weight concrete in the core area. Thermal properties of

normal-weight concrete depend on the type of aggregate. Petrographic inspection by SGH of several

samples of lightweight concrete taken from the debris at NIST showed siliceous sand in the lightweight

concrete. Because source of coarse and fine aggregates is usually the same, the available data for normal-weight concrete with siliceous aggregates are used.

Actual Compressive Strength

Specified concrete strength for lightweight concrete is 3,000 psi and for normal-weight concrete either 3.000 psi or 4,000 psi, as shown on Drawing Book 8, Sheet ABl-2.1 (SHCR 1973). The actual strength,

fg , of in-place concrete at room temperature is calculated from the specified strength, , as follows:

f.-f:-FrF,.F, (1)

where the factor Fi is the ratio of the average strength of cylinders to specified strength. F2 is the ratio of

in-situ 28-day strength to 28-day cylinder strength, and F3 accounts for the change in concrete strength

with age.

By using F, = 1.25 and = 0.95 (Bartlett and MacGregor 1996) and F3 =1.16 based on the formula

specified in Section 2.2.1 of American Concrete Institute (ACI) 209 for change of concrete strength with

age of concrete, the mean of the ratio of actual strength of in-place concrete to the specified concrete

strength I = 1.38. Based on this mean value, the actual strength of in-place concretes are

fg = 5,500 psi for the specified 4,000 psi normal-weight concrete, 4,100 psi for the specified 3,000 psi

normal-weight concrete, and 4,100 psi for the specified 3,000 psi Ughtweight concrete.

Concrete Properties

The unit weight of the lightweight concrete is 100 pcf according to the WTC Design Criteria

(LERA 2003); however. 110 pcf is used based on the two concrete samples described above. The unit

weight of the normal-weight concrete is 150 pcf, according to LERA.

Poisson's ratio, , of 0.17 is used for both normal-weight and lightweight concrete at all temperatures.

Temperature dependent properties of concrete are modulus of elasticity, instantaneous coefficient of

thermal expansion, compressive strength, and tensile strength:

Modulus of elasticity at room temperature is evaluated by the following formula:

EJRT) = 33r;'V^ (2)

The actual strength, , is used as room temperature compressive strength, and 5^[f^ is used as room

temperature tensile strength. Effects of elevated temperature on the listed properties are based on NIST

research (Than 1996, 2003), and plotted in Fig. K-1.

K-5
Appendix K

5-10

410

y 3-10"

(/) J9

UJ

c« 210

1 10

6000

— 5000

4000

Q.E

oo

3000

2000

1000

200 400 Temperature (°C) Normal-weight (SOOOpsi)

Normal-weight (4000psi)

Light-weight (SOOOpsi)

(a) Modulus of elasticity

600

-s.

200 400 600 800 Temperature (°C) Normal-weight

Light-weight

(b) Instantaneous coefficient of thermal expansion

400

^ 300

B 200

200 400 600 Temperature (°C) Normal-weight (SOOOpsi)

Normal-weight (4000psi)

Light-weight (SOOOpsi)

(c) Compressive strength

200 400 600 Temperature (°C) Normal-weight (SOOOpsi)

Normal-weight (4000psi)

Light-weight (SOOOpsi)

(d) Tensile strength

Figure K-1. Temperature-dependent concrete properties.

Concrete Stress-Strain Relationships

The compressive stress-strain curve, based on the formula by Seanz (1964), is given by:

^ s

^

V^c1 J

1 + a + b

£ + c

(3)

K-6
Interim Report on Subsystem Structural Analysis of the WTC Towers

where:

2 c

K.=2, K =

0.85

b = ^-2c , a = c + K^ -2,

K'=^.4^, £^-c1

In tension, stress increases linearly up to the tensile strength. When strained in tension beyond its strength will soften and the stress will drop. However, the descending branch of stress-strain relationship causes significant numerical instability problems which can be avoided by assuming that concrete becomes plastic in tension. Figure K-2 shows a few examples of concrete stress-strain curves at room

and elevated temperatures.

-1000 «

Q.

1^ -2000

0)

-3000

-4000

-0.015

— , /]

*i- X ''I

-4000

-0.01 -0.005

Strain (in/in) T=RT

T=300°C

T=600°C

-0.015 -0.01 -0.005

Strain (in/in) T=RT

T=300°C

T=600°C

(a) Normal-weight concrete (3000psi) (b) Light-weight concrete Figure K-2. Concrete stress-strain curves.

For the knuckle model in LS-DYNA, solid concrete elements are modeled with Pseudo Tensor material

model, where the cap model is used. Since this material model is not temperature dependent, different

material types are specified for the lightweight concrete at RT, 150 °C, 300 °C, 450 "C, 600 °C, and

750 °C (Material IDs 51 through 56) with their different stress-strain relationships.

The concrete slab in the truss is modeled with SHELL181 elements with a concrete material model that accounts for different behaviors in tension and compression. One such material model in ANSYS is the

cast iron plasticity model which uses the Rankine maximum stress criterion in tension, and the expression

for von Mises yield criterion in compression (ANSYS, Inc. 2004). Cast iron plasticity material models

for specified 3,000 psi normal-weight concrete, specified 4,000 psi nonnal-weight concrete, and specified

3,000 psi lightweight concrete are assigned to Material ID 81, 82, and 83, respectively.

K-7
Appendix K

K.5.2 Steel

Steels used in WTC 1 and WTC 2 are listed in Table K-1 along with the yield and tensile strengths used

in our analysis.

Steel Properties

Figure K-3 shows mechanical properties of steel at high temperatures: (a) modulus of elasticity;

(b) Poisson's ratio; (c) yield strength reduction factor; (d) tensile strength reduction factor; and

(e) instantaneous coefficient of thermal expansion. All properties, except yield and tensile strength

reduction factors for bolt steels, are the same for all steels shown in Table K-1.

Stress-Strain Relationship ^

Plasticity: Stress-strain relationships at room temperature were provided by Project 3. They were

constructed from mill report data, actual test data, and literature information using the Voce hardening

law.

Stress-strain relationships at elevated temperatures, without consideration of creep, are obtained by the power law:

a = RrsRcK{T)€ n{T)

ep (4)

where:

K(r) = (/c4-/c0)exp^-0.5

( T \ k2 + [mj [tk2j + kO (5)

n{T) = {n4-nO)exp\-0.5

n1 ( ^

+ 1 [tn^J [tn2j

n2 + nO (6)

The steel stress-strain relationships at different temperatures vary depending on the type of steel used in the construction of the towers. Values for /^v-g, Rq , given in Table K-1, and parameters of K{T) and

n(7) given in Table K-2, were provided by Project 3. The stress-strain curve is linear with Young's

modulus up to the "linearity limif : At the linearity limit, the linear stress-strain curve intersects the power

law stress strain curve. (Stress at the linearity limit is not necessarily equal to the yield stress. The

linearity limit is required for ANSYS input.)

K-8
Interim Report on Subsystem Structural Analysis of the WTC Towers

10

>- 2.5-10

2-10'

0.36

0.34

0.32

0.3

1.5 10 200 400 600 Temperature (°C)

800 1000

0.28

(a) Modulus of elasticity

200 400 600 Temperature (°C)

(b) Poisson's ratio

800 1000

S 0.4

H 0.2

200 400 600 800 1000 Temperature (°C) — Construction Steel

• Bolt Steel

(c) Yield strength reduction factor

200 400 600 800 1000 Temperature (°C) — Construction Steel

• Bolt Steel

(d) Tensile strength reduction factor

200 400 600 800 1000 Temperature (°C)

(e) Instantaneous coefficient of thermal expansion

Figure K-3. Temperature-dependent properties for all steels.

K-9
Appendix K

Table K-1. Steel types used in WTC 1 and WTC 2.

Material ID Description (psi)

OuRT

(psi) RTS RC

1 All 36 ksi core box columns, plates, straps'" 36,720 64,470 1.086 0.857

2 All 36 ksi core WF, channels, and tubes 36 ksi large area and large inertia "rigid" beams in SAP2000 modef

37,000 63,450 1.069 0.954

3 All 42 ksi box columns (1<=0.75 in.) 51,400 79,200 1.070 0.884 4 All 42 ksi box columns (0.75 in. < t <= 1 .5 in.) 47,000 74,800 1.010 0.884

5 All 42 ksi box columns (t > 1.5 in.) 42,600 70,400 0.951 0.880

6 42 ksi or 45 ksi Group 3 WF core coluinns 53,800 74,400 1.005 0.977

7 42 ksi or 45 ksi Group 3 WF core columns 49,000 71,040 0.960 0.954

8 44,200 66 640 0.900 0 948

9 45 ksi Group 4&5 WF core columns 47,800 71,074 0.960 0.939

10 All 36 ksi Plates 1, 2, and 4 in periineter columns 35,630 61,170 1.031 0.875

11 All (42, 45, or 46) ksi Plates 1, 2, and 4 in. perimeter

columns

53,051 74,864 1.011 0.948

12 All 50 ksi Plates 1, 2, and 4 in. perimeter columns. All 50 ksi channels and plates'"

53,991 75,618 1.021 0.978

13 All 55 ksi Plates 1, 2, and 4 with t<=1.5 in. in perimeter

LVJl Ulllllc)

60,817 82,558 1.115 0.903

1 4 /All All AO\j\J L'ci KM Plfitpc r lai^o 1

I tiTiH ^ ix/itn t<'= 1 S in in nprimptpr ^ aiiKi ^ Willi 1 111. Ill uci iiiidCi columns

S7 7sn 1 1 78 0 SQ4

J 5 All AS k'ci Platpc 1 /All \J_' FvjI F ICtlCj 1,

'7 CtllU ?mn id. T Willi f<r=0 l^ \J.^S in111. inIll npnniptpr UCl llllCld columns'*

69 642 90 44'' 1 .22

0\) .y9791 y

16 All 70 ksi Plates 1, 2, and 4 in. perimeter columns 76,735 91,951 1.242 0.955

1/All All 1 ^ l.'s;i IVDl Plf^tPG r JaLCo I ^ nnH zl m nFTimptpr rr^liimnc 1, allU '-r III UCI IlIICLCl CUIUlIIIIo 9.1 469 96 S"* 1 1 308 0\j .y936j\j

1 o All 80 L'ci r\pritTif*tpr pr^liimnc cfpplc rfornrrllpcc r^f" r^lfitp /All OV/ NM UCllIllClCI CUlUllIIIj jlCCIo, ICgdlUlCoD KJl UldlC 91 S 1 7 QO 447 1 343 0\j .yo 987I

1 Q /All All QO inn^ l^ci r^pniTiPtpr rr^liimn ctpplc rfcmrHlpcc \0^, y\J^ l\J\J} IVal UCIIIUCICI CUlUlllII olCCl^, IC^alUlCoo of plate1I11 JS ,"o J 1 . J UVJ \j .y 1 \j

20 Laclede truss web bar rounds specified as A36 38,067 59,567 1.004 0.935

21 Laclede truss chord angels (regardless of ASTM Spec) and all rounds specified as A24255,332 74,050 1.000 0.959

22 A3 2 5 bolts' 104,783 115,983 1.566 0.976

23 All 42 ksi Plate 3 in perimeter columns 42,600 67,216 0.900 0.912 24 All 45 ksi Plate 3 in perimeter columns 45,900 69,831 0.940 0.921

25 All 50 ksi Plate 3 in perimeter columns 51,400 74,188 1.000 0.935 26 All 55 ksi Plate 3 in perimeter columns 56,900 78,546 1.070 0.906

27 All 60 ksi Plate 3 in perimeter columns 62,400 83,903 1.130 0.949

28 All 65 ksi Plate 3 in perimeter columns 67,900 87,261 1.190 0.975

29 All 70 ksi and 75 ksi Plate 3 in perimeter columns 78,900 95,976 1.310 0.997

a. Steels in the following members are assumed to have the properties shown in the table: 36 ksi plates and straps (Material 1 ). 36 ksi channels, tubes, and "rigid" beams (Mateiial 2). 50 ksi channels and plates (Material 12).

b. 65 ksi steels in perimeter columns with t>0.5 in. are assumed to have the same properties as those in Material 15.

c. In the column model, stress-strain relationships of bolts are used. Note: Bolt properties are assumed to be the same as those in Material 19.

K-10
Interim Report on Subsystem Structural Analysis of the WTC Towers

Table K-2. Parameters for k(T) and n(T).

<jyRT = 36,000 psi CT,«r> 36,000 psi

tkl, °C 524.1812 511.8266

tk2, °C 523.6799 511.8938

kO, psi 29049.2 26472.1

kl 9.4346 6.5764

k2 9.3532 6.5971

k4, psi 121605.6 122516.7

tnl, °C 524.4304 519.634

tn2, °C 521.241 499.6031

nO, psi 0.1235 0.0342

nl 19.0000 10.0000

n2 19.0000 10.0000

n4, psi 0.2168 0. 1 5 1

1

Figure K-^ shows stress-strain curves of Material ID 1 (see Table K-1 for the material description) at room and elevated temperatures. Figure K-4 (a) is a close-up view of a low strain range, while

Fig. K-^ (b) shows strain levels up to 0.3.

The elastic-plastic behavior of steels is modeled with ANSYS material model "Multi-linear isotropic

hardening von Mises plasticity."

Elastic + Plastic Strain (in/in) Elastic + Plastic Strain (in/in) T=RT T=RT

T=300°C T=300°C T=500°C T=500''C T=700°C T=700°C

(a) Strain < 0.03 (b) Strain < 0.30

Figure K-4. Stress-strain relationsliips for Material ID 1 steel.

K-ll
Appendix K

Creep: Steel creeps at elevated temperatures (T > 350 °C ), and the creep behavior for steels is based on

the creep model by Fields and Fields (1991), expressed as:

£^^{tJ,a)^^a{T)

where:

V60y

xc(r)

35.5 (7)

yRT J

a{T) =

0

^ Q -(6.1+0.005737)

-| Q-(13. 25-0.008517)

ib(r) = -1.1 + 0.00357

c(r) = 2.1 + 0.00647

for 7 < 350'C

for 350°C < 7 < 500^'C

for 500'^C < 7 < 725X

for 7 < 725"C

for 7 < 725'^C

This model is valid for the temperature range of 350 < 7 < 725 "C . ANSYS uses the "time

hardening creep" model, where creep strain rate is given by:

dt

(8)

and Ci(7), C2(7),and C;^{T) are temperature-dependent parameters determined by Fields' (1991)

creep model given as:

1

' M

a{T)b{Ti

^^'^'^35.5^^'''

60

V <^yR7 ;

' 100

C,(7) = c(7)

C3(7) = ^(7)-1

Figure K-5 illustrates creep behavior of steel at elevated temperatures for Material ID 1 . Figure K-5 (a) shows creep strain rate at different stress levels and different temperatures, and Fig. K-5 (b) compares

elastic, plastic, creep, elastic plus plastic, and total strains at 7 = 400"C and after constant loading for 1,800 s.

Failure Criteria

The failure criteria for steel are defined in terms of plastic strains. The multiaxial fracture strain criterion for different steels and temperatures (Fields 2004) is as follows:

£f = ci'(7)exp

3

2 a

(9)

where stress and strain are true stress and true strain.

K-12
Interim Report on Subsystem Structural Analysis of the WTC Towers

I Plaslic T=500°C Creep T=600°C Elastic Plastic ^— Total

(a) Creep strain rate (b) Comparison of different strains at T=400 °C and t= 1.800 s Figure K-5. Creep behavior at elevated temperatures for Material ID 1 steel.

For the uniaxial stress condition, the plastic strain at fracture reduces to:

^f_uni = exp(-0.5)Qr(r) (10)

Table K-3 shows the uniaxial plastic strain at fracture. £f . at different temperatures calculated by the equation above. This criterion is valid for the fmite element analysis (FEA) with very fine mesh. For

coarse mesh, the equivalent steel fracture criterion was determined numerically as follows. A standard

tension test specunen was modeled in ANSYS. The gauge length, width, and thickness of the specimen were 8 in, 1.5 in, and 1 in., respectively, and Material ID 1 steel properties were used. Six different models (Model 0 to 5) were created, each having a different mesh size. Element sizes of Models 0 to 5 were 0.025 in., 0.050 in., 0.0125 in., 0.250 in., 0.375 in., and 0.75 in. It was assumed that Model 0 was

able to capture tensile fi'acture in a uniaxial tension.

Model 0 was subjected to tension until the maximum plastic strain in the direction of applied

displacement reached the uniaxial fracture strain determined by Eq. (10) for uniaxial stress condition, and

the corresponding elongation of the specimen, Aq, was obtained. Models 1 to 5 were then subjected to the same elongation, Ao, and the maximum plastic strain in the direction of applied displacement was

measured for each model. The maximum plastic strain due to the elongation of Aq is defined as the limiting plastic strain (equivalent fi-acture plastic strain) for the corresponding element size.

From these six cases, a relationship between element size and equivalent uniaxial fracture plastic strain was established. This process was repeated for temperatures 20 °C, 100 °C, 300 °C, 500 °C, and 700 °C.

Figure K-6 (a) shows the ratio of the maximum plastic strain in the direction of applied displacement due

to displacement Aq to uniaxial plastic strain by Eq. (10) vs. element size at different temperatures. The

FEA results were extrapolated up to the element size of 50 in. Plastic strain shown in Fig. K-6 (b) is used

as fi'acture criterion for the corresponding element size in the FEA.

K-13
Appendix K

Table K-3. LIniaxial plastic strain at fracture by Eq. (10).

Plastic Strain at Fracture in the Uniaxial Test Material ID lUU CAA ^AA 1 AAA

1 0.8410.6989 0.6610 1 .0446 1.8100 3.5862

2 0.8410.6989 0.6610 1 .0446 1.8100 3.5862

3 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

4 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

5 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

6 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

7 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

8 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

9 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

10 0.8891 0.7388 0.6987 1.1042 1.9142 3.7907

11 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

12 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

13 0.2846 0.2364 0.2236 0.3534 0.6123 1.2132 14 0.3774 0.3136 0.2965 0.4686 0.8120 1.6088 15 r\ A A /: UA45b A /I 1 AC U.4iy5 U.66zy 1.1456 z.z /5s

16 0.5623 0.4672 0.4418 0.6983 1.2099 2.3972

17 0.7752 0.6442 0.6092 0.9628 1.6681 3.3051

18 0.6545 0.5439 0.5143 0.8129 1.4084 2.7906

19 0.4254 0.3535 0.3343 0.5283 0.9154 1.8137 20 0.8411 0.6989 0.6610 1.0446 1.8100 3.5862

21 0.4908 0.4078 0.3857 0.6095 1.0561 2.0924

=RT tOO'C

=300°C

=500°C

=700°C =RT (power law) MOO'C (power law) :300X (power law) =500°C (power law) :700°C (power law)

—•— EL size = 5.00 in

0.10

(a)

Element Size (In) FEA Results and Extrapolation

400 Temperature (°C) (b) Recommended Limiting Plastic Strain

Figure K-6. Maximum plastic strain from the finite element analysis

and limiting plastic strain.

K-14
Interim Report on Subsystem Structural Analysis of the WTC Towers

K.5.3 Welds

The weld properties at all temperatures are assumed to be the same as those of the base metal of the same

ultimate tensile strength. This assumption is validated by the following observations: the exterior column

welds are strong enough to fail the base metal; the observed fractures in the exterior columns are mostly

through the base metal; and the welds in trusses are resistance welds with no filler added. For the core columns, the area of the welds is significantly less than that of the base metal, and several fractures through the welds have been observed. Fractures in the truss seats and truss connections have also been

observed. High temperature properties of the welding metals have not been found in the literature. Susceptibility of existing cracks in the welds to growth (fracture toughness) does not increase with

temperature (Stevick 1994).

K.5.4 Bolts

A load-elongation relationship for 7/8 in. A325 bolt with 4 in. length at room temperature was provided

by Project 3. Load-elongation relationships at elevated temperatures are constructed by scaling the loads

by the yield and ultimate tensile strength reduction factors for bolt steels shown in Fig. K-3 (c) and (d). Figure K-7 shows the load-elongation relationships of a 7/8 in. bolt at different temperatures. Loadelongation relationships of A325 bolts of different size are scaled by proportioning the load by the ratio of

the bolt thread area to the bolt body area for a 7/8 in. bolt.

0.000 0.050 0.100 0.150 Elongation (in.)

0.200 0.250

Figure K-7. 7/8 In. A325 bolt load-elongation curves at elevated temperatures.

The load-elongation relationship for bolts with a different length than 4.0 in. is expected to be very

similar to the load-elongation relationship of 4.0 in. length as deformations are localized.

Based on the AISC formulas, C-J3-2 to C-J3-4, (AISC 2003), the shear strength for a single shear plane

is calculated as 0.67 of the tensile strength given in Fig. K-7 when threads are excluded from the shear

plane. When threads are not excluded from the shear plane, the nominal shear strength for a single shear

plane is 0.53 of the tensile strength given in Fig. K-7. No shear ductility is assumed at failure.

K-15
Appendix K

K.5.5 Coefficient of Friction

The coefficient of friction of 0.33 for calculation of shear in friction-type connections is the AISC LRFD

(2003) friction coefficient for uncoated clean mill scale steel surfaces, or surfaces with Class A coatings on blast-cleaned steel surfaces.

K.5.6 Symbols

a{T) = temperature-dependent material property that defines fracture criterion

a^{T) = instantaneous coefficient of thermal expansion of concrete

0(^{T) = instantaneous coefficient of thermal expansion of steel

Py{T) = steel yield strength reduction factor due to elevated temperature

P^l{T) = steel ultimate strength reduction factor due to elevated temperature

= unit weight of concrete (110 pcf and 1 50 pcf for lightweight and normal-weight concrete,

respectively)

= Unit weight (490 pcf = 0.284 pci for all steel types at any temperature)

= concrete strain at maximum compressive stress

S^^ = creep strain of steel

£g = elastic strain

= elastic plus plastic strain

£f = effective plastic strain at fracture

uni ^ uniaxial plastic strain at fracture

£p = plastic strain

= concrete strain at maximum tensile strength

£^^J = concrete strain at full crack formation (separation) in tension

= Poisson's ratio of concrete

Vg = Poisson's ratio of steel a = effective stress

<7^ = mean stress

K-16
Interim Report on Subsystem Structural Analysis of the WTC Towers

^yRT ~ room temperature yield strength of steel

^uRT ^ room temperature tensile strength of steel

E^{T) = modulus of elasticity of steel

E^{T) = modulus of elasticity of concrete

F, = mix design factor = ratio of the actual 28 day cylinder strength to

F2 = in-situ factor = ratio of in-situ 28 day strength to the 28 day cylinder strength

F3 = aging factor = ratio of mature concrete strength to 28 day concrete strength

fg = actual strength of in-place concrete

fj = specified 28 day strength

fc{T) = compressive strength of concrete

f({T) = tensile strength of concrete

K{T) = sigmoidal function of temperature with six parameters

n{T) = sigmoidal function of temperature with six parameters

Rq = correction factor that has the following two functions: ( 1 ) to correct the strain rate effect introduced in the material testing and create the stress-strain curve for zero strain rate, and (2) to match the room temperature stress-strain curve at strain of 0.05

Rj^ = ratio of the room temperature tensile strength of the steel of interest to the room

temperature tensile strength of the steel used to develop the power law model

K.6 MODEL CONVERSION FROM SAP TO ANSYS

The SAP2000 (SAP) floor 96 Model of WTC 1 and the SAP Global Model of WTC 1 are converted into ANSYS. The goal of the conversion is to develop ANSYS models that match the SAP baseline models

developed by Project 2 and can be used as a basis of the detailed thermal-structural evaluation. The

converted ANSYS models will be modified to incorporate the nonlinear behaviors of the components and

simplified for the thermal/structural evaluation of collapse initiation study.

K-17
Appendix K

K.6.1 Translation Procedure

Automatic translation software was developed to partially convert the floor model and global model from

SAP2000 to ANSYS 8.0:

• The Joints, Frames, and Shells in the SAP model were translated into ANSYS Keypoints,

Lines, and Areas. Using geometry definition instead of nodes and elements dkectly allows

for ease in local mesh refinement.

• Lines were meshed with both section and real constants so that a translation between

BEAM44 and BEAM188 elements can be achieved by simply changing element types.

Areas were meshed with SHELL63 elements in ANSYS to match the Shell elements in SAP.

Eventually, Lines and Areas will be changed to nonlinear BEAM188 and SHELL181

elements with a type change.

• Material properties were assigned according to the Criteria Document based on the material

definitions and Frame section properties in SAP.

• Frame section properties in SAP were converted into Real Constants for BEAM 44 in ANSYS. Cross section properties in SAP were retained for future conversion into cross

section data for BEAM 188 elements. Shell thicknesses in SAP were converted into Real

Constants for SHELL63 in ANSYS.

• Joint restraints in SAP were translated into DOF constraints in ANSYS.

• Frame distributed loads and area uniform loads were translated into surface loads on Lines

and Areas in ANSYS.

• The ANSYS BEAM44 elements support element moment releases, but the ANSYS nonlinear

BEAM188 elements do not. Therefore, Frame releases in SAP were modeled by coincident

nodes with coupled (CP) degrees of freedom in ANSYS.

• The ANSYS BEAM44 elements allow beam end offsets in three directions, but the ANSYS

nonlinear BEAM188 elements only allow beam end offsets perpendicular to the element axis

through section offset (SECOFFSET) command. Frame insertion points in SAP were

converted in two ways. For offsets along the element axis, additional nodes and rigid

MPC184 elements with the proper lengths were used in ANSYS. For offsets perpendicular to the element axis, beam end offsets were defined using Real Constants for BEAM 44, and

eventually will be defined using SECOFFSET command for nonlinear BEAM 188.

• Frame offsets and rigid panel factor in SAP were modeled by adding additional nodes and

rigid MPC184 elements with the proper lengths in ANSYS.

Those parts of the model that were not converted by the translation software were converted manually.

K-18
Interim Report on Subsystem Structural Analysis of the WTC Towers

K.6.2 Challenges

During the conversion of the SAP Floor Model, the following conditions were encountered and were

resolved:

• The SAP Floor Model allows automatic division of the frames at joints. This causes problems in the translation software because the frame connectivities in the Graphical User

Interface do not show the actual internal element connectivities used in the SAP analysis

engine. In order to resolve this problem, the translation software was modified to use the

internal element connectivities. The table of internal connectivities was exported from the

SAP model after the execution of the analysis.

• Automatic offsets in the SAP model are not available in the ASCII SAP input file prior to the execution of the analysis. The table of element offsets was exported after the execution of the

analysis.

• There are both intentional and unintentional duplicate elements in the SAP Floor Model.

Each leads to problems in the translator since ANSYS cannot have duplicate lines sharing the same key points. Some duplicate elements are used to model additional steel plates at the ends of the trusses. The duplicate elements were manually deleted and the section properties

of the remaining elements were modified to account for the additional steel. Some duplicate

elements are from frame elements which have different lengths and are overlapping each

other. These were manually corrected.

K.6.3 Status

The automatic translation software developed to convert models from SAP2000 to ANSYS was applied

to the floor model and will be applied to the global model shortly. Figures K-8 through K-1 1 show the com erted floor model.

The following analyses were performed to validate the converted ANSYS floor model against the original

SAP model.

• One static analysis with gravity loads as defined in SAP as Load Case "DEAD" which

include self-weight plus 3.5 psf uniform load in the office area.

• One modal analysis, using structural mass only.

Table K-4 summarizes the comparison of the SAP and ANSYS results for the gravity load case. The total reactions for the SAP and ANSYS models are within 0. 1 percent of each other. The maximum slab displacement predicted by the ANSYS model is 3.2 percent smaller than that obtained from the SAP

model. This discrepancy is currently under study and is being resolved. The deformed shapes of the gravity load case for the SAP and ANSYS models are shown in Figs. K-12 and K-13.

Table K-4. Comparison of SAP and A NSYS results or gravity load case. SAP AJSSVS (BEAM 188)

Total reaction, kip 2,212.81 2.210.85 (-0.09 %)

Maximum slab displacement, in. 0.718 0.695 (-3.2 %)

K-1

9
Appendix K

AIM

MAR 22 2004 14:41:28 PLOT NO. 1

Figure K-8. Converted ANSYS model for floor 96 of WTC 1 : overall view.

Figure K-9. Converted ANSYS model for floor 96 of WTC 1 : partial view

near corner of building.

' ELEI'-E>CS I'm m'-i

K-20
Interim Report on Subsystem Structural Analysis of the WTC Towers

Figure K-1 1 . Converted ANSYS model for floor 96 of WTC 1 : view of

floor beams and columns.

K-21
Appendix K

0 .145549 .291097 .436646 .582195 072774 .218323 .363872 .50942 .654969 Figure K-13. Deformed shape (x100) of gravity load case for

ANSYS floor model with BEAM44 (Euler beam) elements.

Table K-5 summarizes the comparison of the SAP and ANSYS resuhs for the modal analysis. The total masses of the SAP and ANSYS models are within 0.02 percent of each other. The dominant natural frequency of the floor predicted by the ANSYS model is 2.5 percent higher than that obtained from the

K-22
Interim Report on Subsystem Structural Analysis of the WTC Towers

SAP model. This discrepancy is consistent with the discrepancy observed for gravity displacement, and

is currently under study and is being resolved. The dominant mode shapes of the floor for SAP and

ANSYS models are shown in Figs. K-14 and K-15.

Table K-5. Comparison of SAP and ANSYS Modal Analysis Results.

SAP ANSYS (BEAM 188)

Total mass, Ib-sec'/in. 5448.7 5447.7 (-0.018%)

Dominant natural frequency of floor, Hz 4.32 4.43 (+2.5 %)

Figure K-14. Donriinant mode shape (frequency = 4.32 Hz) of floor structure

for SAP floor model.

Msdal analysis, TlTOshenko bea-rs (BEaC83) - quadratic shape function

Figure K-15. Dominant mode sliape (frequency = 4.43 Hz) of floor structure

for ANSYS floor model.

K-23
Appendix K

K.7 FULL FLOOR SUBSYSTEMS

The full floor model is analyzed using the ANSYS general purpose finite element program Version 8.0. The objectives of the analysis are:

• To identify the most likely failure modes,

• To evaluate - Failure loads, - Temperatures at failure, - Time-to-failure, and r , - Changes in mechanical properties and geometry at failure.

• To simplify the model and to reduce the computational efforts for incorporation into the

global model.

The failure modes and the failure loads of different components of the full floor subsystem are evaluated

through analysis of detailed models of those components, using either hand calculations or FEAs.

Simplified models that capture the failure loads and failure modes are then developed for each

component. These simplified models of components are incorporated in the full floor subsystem model.

In this chapter, after a general description of the full floor model, the analyses of important components

are presented and discussed.

K.7.1 Full Floor Model

Model Description

The floor model is developed using the converted SAP2000 model for floor 96, with the following

modifications:

1. Combine two adjacent trusses into a single truss. The elements in the truss model have

double the areas of elements in each real truss.

2. Change rigid beams at knuckle locations to user-defined elements with the properties of the

knuckle determined by the component knuckle model.

3. For compression diagonals, add user-defined elements to account for buckling of diagonals.

4. For truss ends and connections, add user-defined elements to account for truss seat failure.

5. Pin concrete slab for out-of-plane rotation at both its interior and exterior edges.

6. Use user-defined elements along the edge nodes of the concrete slab to model the tensile strength of the concrete slab and the in-plane shear capacity at the connection to the spandrel

plate.

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Interim Report on Subsystem Structural Analysis of the WTC Towers

7. Remove the spandrels defined as beam sections in SAP2000 model and replace them with

SHELL181 elements in ANSYS. (This modification eliminates the need for defining panel zone stiffness.) The new spandrels will wrap continuously around the floor. Each spandrel

plate between columns will be represented by 16 elements, 4 in. height and 4 in. width.

Material and geometry assignments are carried through to ANSYS.

8. Change the elastic column elements as translated into ANSYS to user-defined sections with

BEAM189 elements with plasticity and creep.

9. Make new column sections within the limits of the spandrels with reduced Plate 3 thickness,

say 0.005 in. in thickness, to insure correct modeling of torsional stiffness. Spandrel

thicknesses should be reduced within the limits of the column by the same thickness.

Connect the centerline of column to spandrel with rigid elements.

Material Properties

ANSYS's multilinear isotropic hardening von Mises plasticity with time hardening for temperatures

below 350 °C is used for the beam elements representing the truss system, girders, beams, and columns in areas where plasticity is either anticipated or expected to occur by analysis. This material model with

creep is used for temperatures above 350 °C. This material model is used for shell elements representing

the spandrel plates, when appropriate.

Loading

The full floor model is analyzed for dead and live loads first, and then thermal loads are applied to model

the path dependent nonlinear response. The thermal loads are provided by Project 5 and include

temperatures and temperature gradient time-histories for all structural members in the full floor model for

(1) standard fire, (2) representative building fire scenarios, and (3) different fire protection scenarios.

Boundary Conditions

The beam elements representing the columns are restrained vertically at floor 95. The outward and

tangential displacements and all rotations of the column ends at floors 95 and 97 are fixed to restrain thermal expansion. Mass elements defined by the tributary dead and live loads are added to the top of the columns and at connections to floor 96.

Failure Modes

The possible failure modes of the floor subsystem are as follows:

1 . Sagging: Floor sagging along the axis of the main trusses may be caused by ( 1 ) loss of

stiffness and softening of truss at high temperature, (2) catenary action of the truss due to

plastic bending or buckling of critical members required for truss action, or (3) loss of

composite action of floor-to-knuckle failure. These are discussed in some detail under truss

failure modes. Floor sagging may result in component failure due to tension in the truss

seats, tension in the floor subsystem, tension on the connections to the exterior walls, lateral

K-25
Appendix K

loads on columns, and increased demand on other components of the floor subsystem, for example, bridging trusses and transfer trusses and their connections.

2. Edge Sagging: Edge sagging results from failure of truss seat connections at either the

interior or exterior supports and is evidenced in videos. Edge sagging, similar to sagging,

increases demand on other components of the floor subsystem, reduces buckling strength of

columns, and can lead to failure of a floor.

3. Loss of Support: Abrupt failure of the floor subsystem can result from loss of truss support

for a large number of adjacent parallel trusses. Loss of a truss support can occur due to

( 1 ) vertical shear load due to debris and/or impact load of the dropping floor above,

(2) vertical and horizontal shear loads resulting from slab expansion acting on column truss seats (3) tension acting on column truss seats, and (4) cooling of a truss shortened by plastic

deformation and loss of composite action. Failure of truss support will increase the demand

on the adjacent trusses and can result in sequence of truss seat failure, edge sagging, and

ultimately failure of the floor subsystem.

4. Expansion of Floor System: Expansion of floor resuhs in deformation of columns and forces

at comers of the exterior wall subsystem. Such comer forces can initiate a failure sequence

of columns near the comers. Such a failure includes development of horizontal shear in the

gusset plates and the exterior column tmss seats, large forces in the straps, and large lateral x and y forces in columns, especially near the comers.

K.7.2 Knuckle Analysis

The "knuckle" is formed by the extension of the tmss diagonals into the concrete slab and provides for composite action of the steel truss and concrete slab. The composite action is due to the shear transfer between the knuckle and the concrete slab both in the truss transverse and longitudinal directions.

The objective of this analysis is to predict the knuckle capacity when the tmss and concrete deck act as a composite member and to develop a simplified model of the knuckle behavior to be included in the full floor subsystem model.

Knuckle Shear Tests

Two sets of experiments were performed in 1967 at Laclede Steel Company in Saint Louis, Missouri, to determine the transverse and longitudinal shear capacities of the knuckle.

The transverse shear test consists of double knuckles placed into two reinforced concrete blocks that were

confined on the comers by angles as shown in Fig. K 16. The concrete density of 1 10 pcf corresponds to the lightweight concrete in the office areas. The concrete compressive strengths reported for 7 day and

27 day cylinder tests were 1,330 psi, and 2,600 psi, respectively. The inner ends of the two knuckles were

connected through channels to a #1 1 rebar and the rebar was loaded until the concrete failed. The tests were conducted at concrete age of 6 and 27 days. The primary failure mode observed was concrete shear

failure. The pictures from the tests show formation of the shear crack in one of the concrete blocks and

edging of the channel into the concrete. The transverse shear capacity of the knuckle as the average of the

K-26
Interim Report on Subsystem Structural Analysis of the WTC Towers

yfCT/OA/ A-A Drawing provided by Laclede Steel. Figure K-16. Transverse shear test of a knuckle.

two reported tests is 16.9 kip per knuckle. After adjusting it for the strength of in-place, mature,

lightweight concrete in the slab of 4,100 psi relative to the average strength of the lightweight concrete used in the test of 1,965 psi, by multiplying by the ratio of 4,100 to 1,965 psi, the transverse shear

capacity of the knuckle is approximately 35 kip per knuckle.

The longitudinal shear test consists of double knuckles placed into two concrete blocks as shown in Fig. K-17. The test specification shows comer angles confining concrete blocks and no reinforcement for the concrete. However, the test pictures show reinforcement in both directions for each concrete block,

with the comer angles dismantled. The test specification calls for concrete density of 152 pcf, which

corresponds to a normal-weight concrete. The slab in office areas is of lightweight concrete. The average

strength of two 28 day cylinders tested is 4,290 psi. A third sample, tested after 96 days, showed a strength of 2,850 psi. The test specification does not identify the weld size connecting the inner ends of

the two knuckles to two channels. However, the primary failure mode observed for three tests is weld

failure. Weld failure is not identified as the failure mode of the knuckle for the other two tests. The

results of the shear tests of the knuckle in the longitudinal direction based only on these two tests are approximately 28.3 kip per knuckle. After adjusting for the strength of in-place, mature, lightweight concrete in the slab of 4,100 psi relative to the 28 day corrected average strength of the normal-weight

concrete used in the test of 3,707 psi, by multiplying by the ratio of 4,100 psi to 3,707 psi, the longitudinal shear capacity of the knuckle is approximately 3 1 kip per knuckle.

Drawing provided by Laclede Steel. Figure K-17. Longitudinal shear test of a knuckle.

K-27
Appendix K

The effect of temperature on the knuckle is as follows:

• The steel knuckle conducts the temperature of the diagonal without much loss into the cool

concrete. Concrete has a low coefficient of conductivity and does not respond rapidly to the

rise of temperature.

• Concrete in the intermediate proximity of the metal knuckle will heat to a temperature close

to that of the steel.

• Shear failure of the knuckle is initiated by the failure of concrete in close proximity to the

knuckle. Final failure will engage not only the hot concrete in close proximity of knuckle,

but the cooler concrete farther away.

• It is reasonable to assume that for gas temperatures in the range of RT to 450 °C, 650 °C,

850 °C, and 1050 X , the knuckle metal temperature is below 375 °C, 550 "C, 725 "C, and

900 °C, and the average concrete temperature is below 300 "C, 450 °C, 600 °C, and 750 °C,

respectively.

Neglecting the difference in thermal expansion of concrete and steel, for gas temperatures of RT, 450 °C,

650 °C, 850 °C , and 1050 °C, the expected concrete strength is in the range of 4,100 psi, 3,300 psi,

2,600 psi, and 2,000 psi, and the knuckle capacity in either direction is 30 kip, 24 kip, 19 kip, and 15 kip,

respectively.

Knuckle Test Finite Element Model

Finite element models, shown in Fig. K-18, represent one quarter of the knuckle test specimens. The

knuckle and channel members in the test set up are modeled by solid steel elements. Concrete Pseudo

Tensor model and the LS-DYNA computer program were used for the analysis. An imposed ramped

displacement was applied to the angle member.

The concrete strength used in the finite element model for the longitudinal shear of the knuckle was

4,100 psi and for the transverse shear of the knuckle was 2,500 psi. In addition 0.47 percent steel reinforcement representing welded wire fabric reinforcement of the slab was added in a distributed way to the concrete. Also, two different assumptions were made about the interface condition between the concrete and the steel: fully bonded and frictionless.

The results of the analysis are shown in Figs. K-19 through K 22. They show significant dependence on

the characteristic of the interface between the steel and concrete. The longitudinal shear test FEA results,

shown in Fig. K-21, show that each knuckle has strength in the range of 15 kip to 35 kip, depending on

interface. The test results show that the interface is closer to fully bonded case. For the transverse shear,

the FEM results. Fig. K-22, show that transverse knuckle strength is about 24 kip for 2,500 psi concrete,

corresponding to 39 kip for 4,100 psi concrete. Figure K-20 shows that for transverse shear concrete crushes in a small region next to the knuckle and extending in front of the shear load. Figure K-20 also

shows large regions of crushing at the lower boundary of the model. These regions are the result of

imposing the boundary condition UY=0. This boundary condition, and the crushing at the boundary,

although realistic for the test, would not obtain in a pair of transversely loaded knuckles of the double

K-28
Interim Report on Subsystem Structural Analysis of the WTC Towers

(2.1 In. region)

Figure K-18. Finite element models of knuckle shear tests.

HODAL SOLUTION

SI2P-1 SUB -27 TIME-. 13

E3 (iVG) DMK -.077SB1

SMH —10872

SMX -136.423

L

1

AN

APR 16 2004 13:37:01

-4100 -3069 ^^^^^BmS 1 -1067 -55.556 -3594 -1572 J -561.111 450 Shear Parallel to Truss. ii i^HWIl^li ' (psi)

Figure K-19. Compressive stresses in longitudinal shear finite element model.

K-29
Appendix K

-2400 Shear Perpendicula-1111 -466.667 177.778

.869 -144.444 500 Stress in Coirrrete (psl)

Figure K-20. Compressive stresses in

transverse shear finite element model.

80000

0 • ' ' 1 0.00 0.02 0.04 0.06 0.08 0.10 0.12

Displacement (in.)

Figure K-21. Shear force versus displacement from finite element

model for longitudinal shear of two knuckles.

50000

45000

40000

35000 J

30000

J 25000 ^ 20000

15000

10000

5000

0.00

0 Friction Contact

Fully Bonded

0.05 0.10 0.15

Displacement (in.)

0.20 0.25 0.30

Figure K-22. Shear force versus displacement from finite

element model for transverse shear of two knuckles.

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Interim Report on Subsystem Structural Analysis of the WTC Towers

truss. The small crushing regions at the knuckle indicate that a pair of knuckles in a double truss can be

expected to behave nearly independently of each other, and, therefore, have nearly double the capacity of

a single knuckle. Unfortunately, test results are not available that would confirm this conclusion.

Although the analysis shows the sensitivity of the results to the interface assumptions, it justifies the shear

capacities computed from the test results.

Knuckle Model

The purpose of the detailed finite element analysis of the knuckle is to provide a basis for deriving a simple model for use in analyses of the full floor. The knuckle model includes segments of concrete floor and truss diagonal that protrudes into the 4 in. thick concrete. The dimensions of the concrete included in the model are one half of the double truss spacing of 40 in. The diameter of the truss diagonal included in the model is 1.09 in., and the center line of the knuckle is 3 9/16 in. from the center line of the double

truss. The concrete slab wire mesh reinforcement is modeled by distributed reinforcement properties.

The model is boimded by four planes. Two of these planes are parallel to the chord of the truss, and the other two planes are perpendicular to the chord. Symmetry conditions are applied to these planes

consistent with the loads. For the tension loading, in addition to the symmetry conditions, the model is supported vertically at both symmetry planes that are perpendicular to the truss chord. For the shear load

parallel to the chord, the model is supported ahead of the shear load in the direction parallel to the chord.

The knuckle has the properties of ASTM A36 (Material ID 20) round bar steels and the concrete has

lightweight concrete properties specified for LS-DYNA with concrete-cap model.

K.7.3 Column Truss Seats

In this section, likely failure modes of truss seats are identified, and the corresponding failure loads are determined. The following loading conditions were considered: vertical force, horizontal tensile force,

horizontal compressive force, and combined vertical and horizontal force.

Description of Column Truss Seats

The floor truss is supported at the exterior wall and at the core by seats. The truss seat at the exterior wall

and at the core will be referred to as exterior seat and interior seat, respectively.

The interior seat consists of a horizontal plate with two vertical plate stiffeners as shown in Fig. K-23.

These plates are fillet welded together and fillet welded to the core channel beam. Two 5/8 A325 bolts (one bolt in each truss) connect the truss to the seat. The bolt connection is a friction type connection

with 1 3/4 in. long slotted holes in the seat and 7/8 in. oversize holes in the bearing angles.

K-31
Appendix K

Side View Front View Top View

Figure K-23. Interior seat.

The exterior seat consists of a seat angle attached to the spandrel with two vertical plates (stand-off

plates), and a gusset plate as shown in Fig. K-24. Fillet welds connect the seat angle to the stand-off, the

stand-off to the column/spandrel, and the gusset plate to truss top chord. A complete-joint-penetration

groove weld connects the gusset plate to the column/spandrel. Similar to the interior seat, each pair of

trusses is attached to the exterior seat by two 5/8 in. A325 bolts. The bolt connection is a friction type

connection with 2 in. long slotted holes in the seat angle and 7/8 in. oversize holes in the truss-bearing

angle.

Side View Front View Top View

Figure K-24. Exterior seat.

In floor 96 of WTC 1 , there are seven types of interior seats and eight types of exterior seats. The

different types of interior seats are identified with Detail Numbers 15, 17, 20, 21, 22, 23, and 226A; and

the exterior seats with Detail Numbers 1013, 1111, 1212, 1311, 1313, 1411, 151 1, and 1611, as shown in Fig. K-25.

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Interim Report on Subsystem Structural Analysis of the WTC Towers

Original drawing provided with permission from PANYNJ.

Figure K-25. Truss seat detail location on northeast quadrant of floor 96 of WTC 1.

All types of interior seats are similar in their design, but are all unique because of the variation in the size of the vertical and horizontal plates, the location of the bolt holes, and the size of the fillet welds. The

thickness of the plates ranges from 0.375 in. to 0.75 in.; the distance between bolt holes ranges from

8.5 in. to 10.5 in.; and the size of the fillet welds ranges from 0.25 in. to 0.375 in. All types of exterior

seats are also similar in their design, but are all unique because of the variation in the size of the stand-off,

and size of the seat angle, the size and shape of the gusset plate, the location of the bolt holes, and the size of fillet welds. The vertical height of the stand-off ranges from 8 in. to 1 1 in. The smallest seat angle size

is L4 X 4 X 1/2, and the largest is L6 x 4 x 3/4. The shapes of the gusset plate are rectangular and

trapezoid, and the plate ranges in width from 4.5 in. to 6 in. The distance between bolt holes ranges from

3.25 in. to 10.5 in., where it is 3.25 in. when the seat is supporting a single truss. The size of the fillet welds ranges from 0.2125 in. to 0.375 in.

Truss Seat Material Properties

The material properties used in the calculations were selected from Table K-1 to best match the material

properties indicated in the design drawings. Figure K-3 was used to determine the material mechanical

K-33
Appendix K

properties at high temperature. The material properties used for truss seat calculations are summarized in Table K- 6.

Table K-6. Material properties used for truss seat calculations.

Description Selected Material ID

Exterior and interior seat A325 bolts Material 22

Fillet welds Material 7 Truss bearing angles Material 2

Exterior seat Seat angle Material 1 Gusset plate Material 12 Stand-off Material 23

Truss top chord angles Material 21 Cover plate for bridging truss top chord Material 1

Interior seat Vertical plate stiffener Material 12

Horizontal plate Material 12

Truss Seat Failure Modes and Sequence

The failure modes of different truss seats are identified for vertical force, horizontal tensile force,

horizontal compressive force, and combined vertical and horizontal force.

Failure Modes ofInterior Seat against Vertical Force: The location of the vertical load on the truss seat

is eccentric to the plane of fillet weld connection between the truss seat and the channel beam. Hand

calculations have shown that the fillet welds at this connection, which must resist shear and bending,

control the truss seat capacity. The failure mode is fracture of the fillet welds at this connection, which

results in loss of the truss vertical support.

Failure Modes ofInterior Seat against Horizontal Tensile Force: The failure modes considered are

(1 ) bolt shearing, (2) bolt bearing, (3) bolt tear-out, and (4) block shear failure. Hand calculations have

shown that the bolt shear strength controls the truss seat capacity. Bolt shear by itself, however, does not cause the truss to lose its vertical support, but it is the prerequisite to the truss walking off the seat. The

travel distance required for the truss to walk off of the seat is 4 in.

Failure Modes ofInterior Seat against Horizontal Compressive Force: The concrete slab above the

truss seat connection provides the compressive force resistance. If the concrete slab fails, the truss seat has resistance against compressive force from bolt fi^iction and surface friction between the seat and

bearing angles. Additional resistance is developed when the truss comes into contact with the channel

beam. Travel distance for the truss to come into contact with the channel beam is 1/2 in. Under

compressive force, the truss will not lose its vertical support.

Failure Modes ofInterior Seat against Combined Vertical and Horizontal Forces: Under combined

vertical and horizontal forces, the failure modes are a combination of the individual failure modes for

vertical and horizontal forces.

Failure Modes ofExterior Seat against Vertical Force: The location of the vertical load on the seat is eccentric to the plane of connection between the seat and the spandrel. Because of this eccentricity, the

K-34
Interim Report on Subsystem Structural Analysis of the WTC Towers

truss seat must resist both shear and bending. Finite element analysis of the truss seat was used to determine load paths and evaluate the behavior of the seat connection.

Figure K-26 shows the finite element model of the seat connection, where half of the seat was modeled

and symmetry boundary conditions were applied. The results of the finite element analysis show that shear force is carried primarily by the stand-off plates shown in Fig. K-24, while the bending moment is resisted by tensile force in the gusset plate and compressive force in the stand-off plate. The seat restrains the moment until horizontal force in the connection causes slippage between the seat angle and bearing

angle. Fillet welds at the stand-off to spandrel connection, which must resist shear, bending, and

compression, control the seat capacity. The failure mode is fracture of the fillet welds as this connection,

which results in loss of truss vertical support.

Failure Modes ofExterior Seat against Horizontal Tensile Force: The failure modes considered are:

(1) failure of the groove weld between gusset plate and spandrel, (2) failure of the fillet weld between the gusset plate and the truss top chord, (3) tensile failure of the gusset plate, (4) bolt shearing off, (5) bolt

bearing, (6) bolt tear-out, and (7) block shear failure. For calculation purposes, the bolts are assumed to be centered in the slotted holes. The typical failure sequence of the truss seat is as follows: first the gusset

plate yields, then it fractures, followed by truss deformation and boh bearing against the slotted hole, the

bolt shears off, and then finally the truss walks off the seat. The travel distance for the truss to walk off of

the seat is 4 5/8 in. This failure sequence is illustrated in Fig. K-27 as path (A) and shown in Fig. K-28,

where the relationship between the tensile force resistance from the seat connecfion and the truss travel distance is plotted. In this plot, frictional resistance between the seat angle and bearing angle was not

included.

Gusset plate

Apply load and ^ boundary conditions

at diagonal strut

Figure K-26. Finite element model of exterior seat.

K-35
Appendix K

Gusset plate

yields

(A)

Gusset plate

fractures

(B) Fillet weld

fractures

Bolt

shears off

(C) Bolt

shears off

Fillet weld

fractures

Truss walk-off seat

(A) Seat details 1111, 1311, 1411, 1511, and 1611 at all temperatures.

(B) Seat detail 1013 at temperatures below lOOT.

(C) Seat details 1212 and 1313 at all temperatures, and detail 1013 at temperatures more than or equal to 100 °C. Figure K-27. Failure sequence of the exterior seats against tensile force.

160

C

Gusset plate yields Gusset plate fractures

700"C

At travel distance 4-5/8 in. truss walks off support

Bolt shears off

Slip resistance from

bolt connection —

—I

0.4

20°c -

300°C 400°C

500°C

600°C

700°C

2 )0'

0.2 0.4 0.6

Truss travel distance (in.)

0.8

Figure K-28. Typical tensile force resistance from exterior seat (Detail 1411).

Seat details 1212 and 1313 have a wider gusset plate and follow path (C) which differs from the typical

sequence where the bolts will bear against the slotted hole then shear off before the gusset plate

connection fails. The failure sequence of seat detail 1013 is temperature-dependent. At temperatures

below 100 "C, the fillet weld connection between the gusset plate and the truss top chord fractures before

bolts shear off. At temperatures greater than or equal to 100°C the failure sequence is the same as for Details 1212 and 1313.

Failure Modes ofExterior Seat against Horizontal Compressive Force: The concrete slab above the

truss seat connection provides the compressive force resistance. If the concrete slab fails, the truss seat

K-36
Interim Report on Subsystem Structural Analysis of the WTC Towers

has resistance against compressive force provided by the gusset plate until it buckles, and from bolt

friction and bolt shear until the bolt bears against the slotted hole and then shears off Surface friction between the seat angle and bearing angles will also provide some resistance. Additional resistance is developed when the truss comes into contact with the spandrel. Travel distance for the truss to come into contact with the column spandrel is 1 1/2 in. Under compressive force, the truss will not lose its vertical support.

Failure Modes ofExterior Seat against Combined Vertical and Horizontal Force: Under combined

vertical and horizontal forces, the failure modes are a combination of the individual failure modes for

vertical and horizontal forces.

Truss Seat Capacity Calculations

In this section, truss seat capacities corresponding to the failure modes described in the previous section

are given. The capacity is computed for the different types of the truss seat at different temperatures.

Calculation of the connection capacity was performed using the methods in the Manual ofSteel

Construction: Load and Resistance Factor Design (AISC 2001) with the resistance factor, (p, assumed to be equal to one.

Capacity ofInterior Seat against Vertical Force: Failure mode of the truss seat against vertical force is fracture of the fillet welds at the seat-to-channel beam cormection. Strengths of the fillet welds at this connection are summarized in Table K-7. The symbol # in this table refers to seat detail number.

Table K-7. Interior seat capacity against vertical force.

Temp.

(°C)

Connection Capacity Against Vertical Force (kip)

#15 #17 #20 #21 #22 #23 #226A

20 226 226 265 221 187 187 385

50 226 226 265 221 187 187 385

100 226 226 265 221 187 187 384

200 225 225 264 220 187 187 383

300 220 220 258 215 182 182 374

400 201 201 236 197 167 167 343

500 160 160 188 156 132 132 272

600 98 98 116 96 82 82 167

700 45 45 53 44 37 37 76

Capacity ofInterior Seat against Horizontal Tensile Force: Failure loads were computed for the failure modes described above. Table K-8 summarizes the results for Seat Detail 22. This table shows that the shear strength of the two bohs controls the horizontal tensile strength of the truss seat connection. As can be seen from this table at temperature 500 °C, bolt shear capacity is reduced by half, and at 600 °C it is reduced to less than a quarter of the original capacity at room temperature. Other seat details also have

the same failure mode, and, therefore, the same failure load.

K-37
Appendix K

Table K-8. Interior seat capacity against tensile force.

Temp.

(°C)

Resistance against Tensile Force (kip)

Bolt Slip Critical

Bolt Shearing Off

Bolt Bearing Bolt Tear-out Block Shear

On Seat On Truss From Seat From Truss Of Seat Of Truss

20 6 44 124 69 87 101 60 59

50 44 124 69 87 101 60 59

100 6 44 124 69 87 101 60 59

200 6 44 124 69 87 100 59 59

300 6 42 121 68 85 98 58 57

400 6 34 111 62 77 90 53 52

500 6 21 88 49 61 71 42 42

600 6 9 54 30 38 44 24 24

700 6 4 25 14 17 20 10 10

Capacity ofInterior Seat against Horizontal Compressive Force: Under compressive force, the truss will come into contact with the channel beam before the boh bears against the slotted hole. The truss seat connection does not fail under compressive force.

Capacity ofInterior Seat against Combined Vertical and Horizontal Force: A typical interaction

relationship for combined vertical and horizontal tensile force is shown in Fig. K-29. As can be seen from this figure, the vertical shear strength of the seat reduces because of the additional horizontal tensile force that the fillet weld connection between the truss seat and the channel beam must resist.

200

0 '

^ ^ r-i 1 r-^ 1

0 10 20 30 40 50

Horizontal tensile strength (kip)

Figure K-29. Strength of combined vertical and horizontal force (Detail 22).

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Interim Report on Subsystem Structural Analysis of the WTC Towers

Capacity ofExterior Seat against Vertical Force: The failure mode of the truss seat against vertical force

is fracture of the fillet welds at the stand-off-to-spandrel connection. Strength of the fillet welds at this connection is summarized in Table K-9.

Tab e K-9. Exterior seat capacity against vertical force.

Temp.

(°C)

Connection^Capacity against Vertical Force (kip)

#1013 #1111 #1212 #1311 #1313 #1411 #1511 #1611

20 94 94 111 94 94 140 193 207

50 94 94 111 94 94 140 193 207

100 94 94 111 94 94 140 193 207

200 93 93 110 93 93 139 192 206

300 91 91 108 91 91 136 187 201

400 84 84 100 84 84 126 172 184

500 69 69 81 69 69 102 136 146 600 45 58 53 60 45 78 84 90

700 29 26 34 27 29 35 38 41

Capacity ofExterior Seat against Horizontal Tensile Force: The connection capacity of truss seats that follow failure sequence (A), as shown in Fig. K-27, equals the failure load for mode (3) defined

previously. The connection capacity of truss seats that follow failure sequence (B) equals the failure load

for mode (2). The connection capacity of truss seats that follow failure sequence (C) equals the failure load for mode (4) plus the developed resistance from the gusset plate. The results of the exterior seat capacity calculations are summarized in Table K-10. Note that the strength of the truss seat #1013

increases by about 38 percent at a temperature of about 100 °C. For temperatures less than 100 °C, the capacity is controlled by the gusset seat fillet weld strength, and for temperatures in excess of 100 °C, the

bolt reaches the end of its travel in the elongated bolt hole and increases the capacity of the connection.

Table K-10. Exterior seat capacity against

horizontal tensile force.

Temp.

(°C)

Connection Capacity against Tensile Force (kip)

#1013 #1111 #1212 #1311 #1313 #1411 #1511 #1611

20 100 104 182 134 182 134 134 134

50 100 104 182 134 182 134 134 134

100 138 104 181 134 181 134 134 134

200 135 103 180 133 180 133 133 133 300 130 101 174 130 174 130 130 130

400 115 93 156 120 156 120 120 120

500 84 75 117 96 117 96 96 96

600 42 49 67 62 67 62 62 62

700 20 25 32 31 32 31 31 31

Capacity ofExterior Seat against Horizontal Compressive Force: Under compressive force, the gusset

plate will buckle before the bolts shear off Compression strength of the gusset plate governs the truss

seat capacity. The compressive strength of the gusset plate is summarized in Table K-1 1.

K-39
Appendix K

Table K-11. Compression strength of gusset plate.

Temp.

(C)

Compression Strength of Gusset Plate (kip)

#1013 #1111 #1212 #1311 #1313 #1411 #1511 #1611

20 77 69 99 90 99 90 90 90

50 76 68 98 89 98 89 89 89

100 74 uu o / Qf, yyj 87 87 87

200 71 63 91 83 91 83 83 83

300 67 60 87 79 87 79 79 79

400 62 55 80 72 80 72 72 72 500 48 42 61 55 61 55 55 55 600 20 17 25 22 25 22 22 22

700 6 5 8 7 8 7 7 7

Capacity against Combined Vertical and Horizontal Force: Interaction relationships for combined

vertical and horizontal forces are under development.

K.7.4 Modeling Connection Failure by Break Elements

In this section, simplified finite element models of the exterior and interior seat, knuckle, stud on strap

anchor, and stud on spandrel are described. These connection models were developed for incorporation

in the floor truss analysis to capture the connection failure modes and deteraiine the sequence of the

failure modes that may lead to the failure of the floor truss.

The developed simplified model of these connections simulates the loss of connection resistance after failure either by exceeding the connection force capacity or by exceeding the allowable deformation (truss

walking off the seat). The connection capacity can also be temperature-dependent. The finite element

modeling assumptions are as follows:

Break element, a unidirectional linear spring element with the capability of turning on and off during an

analysis, is used for modeling connection failure. The element is a part of the structure that connects two

"active" nodes in the "on" mode and disconnects them in the "off mode, depending on the relative displacement of two "control" nodes. The break element is defined as follows:

B„\_iI,j,dof,j);(k,l,dofki);(K,Aon (H)

where m is the break element number, / andj are the active nodes, dofij is the degree of freedom for the

active nodes, k and / are the control nodes, dofki is the degree of freedom for the control nodes, K is the

elastic stiffness of the break element, and Aq is the differential displacement limit of the control nodes.

A beam element with temperature-dependent thermal expansion material properties is used to make the

connection capacity temperature-dependent. This is done by using the deformation of the beam element

from thermal expansion to control the status (on/off) of the break element. Figure K-30 illustrates the basic mathematical model of the connection. The connection capacity is made temperature-dependent by

defining the thermal expansion of the beam element to be temperature dependent.

K^O
Interim Report on Subsystem Structural Analysis of the WTC Towers

Basic mathematical model:

Let node / and / have same displacement.

- In this example, control node k and / relative displacement has the form:

U, -U, = — + oLbJ

K

- Control element status is by definition

U,^ - U, < , status : on

U,. -Ui > Aq, status : off

- Define break element to turn off when capacity, P„. has been reached

Ao = —+ C(LAT K p

- Element capacity at room temperamre is given as

- Hence for a given P„ and AT relationship, the required thermal expansion to turn element off can be found by

P

J k

Control element Beam element

- Stiffness. K - Length. L

aPluin -PII Km

Figure K-30. Basic mathematical model of connection failure.

Multiple connection failure modes require use of different break elements that are connected together in a

logical manner. For example, to model independent failure modes, that is, one failure mode that does not cause other failures, break elements are connected in parallel. If one break element turns off, the other

break elements remain. For dependent failure modes, break elements are connected in series. If one

break element turns off, then all elements turn off.

Simplified Model of the Interior and Exterior Seat

The failure modes of the interior seat include ( 1

) the truss walking off the support, (2) exceeding the

vertical temperature-dependent shear capacity of the truss seat, and (3) exceeding bolt temperature-dependent shear capacity when bolt bears against slotted hole. These failure modes are

K-41
Appendix K

captured by using four break elements and two beam elements as shown in Fig. K-3 1 . Results of the

simplified seat model capturing failure from the truss walking off support and failure from exceeding seat

vertical shear capacity are shown in Fig. K-32 and Fig. K-33, respectively, which depict the relationship

between the horizontal and vertical seat forces and the horizontal truss travel distance.

When truss reaction force on the seat is large in horizontal tension and small in vertical shear, the failure mode is bolt shearing off followed by truss walking off the support as shown in Fig. K-32. Bolt shear is controlling the seat horizontal resistance capacity. Bolt shear by itself however does not cause the truss to lose its vertical support, but it is the prerequisite of truss walking off the seat. The travel distance for a

truss to walk off an interior seat is 4 in. When truss reaction force on the seat is large in vertical shear and

small in horizontal tension, the failure mode is exceeding the seat vertical shear capacity as shown in Fig. K-33. This failure mode will cause the truss to lose both its vertical and horizontal support from the

seat. ' \

The simplified model of the exterior seat is the same as the simplified model of the interior seat, except

for an additional beam element and a break element to model failure of the gusset plate shown in Fig. K-34.

Simplified Model of the Knuckle

Knuckle failure modes that must be captured by the simplified model are the horizontal shear and vertical

tensile failure, which are both temperature-dependent. Finite element modeling assumptions for the

knuckle are: (1) the knuckle has resistance in all translational DOF, (2) the knuckle does not have a

vertical compression capacity limit, (3) capacities in the horizontal shear and vertical tension are dependent, and (4) vertical compression resistance is independent of the capacities in the other directions.

Knuckle failure is captured by using 15 control elements and 5 beam elements as shown in Fig. K-35.

Simplified Model of the Stud on Strap Anchor and Stud on Spandrel

Simplified models of stud on strap anchor and stud on spandrel were developed using the same technique

as described for the knuckle model.

K.7.5 Truss Model

Objectives

The objectives of the truss model study are to:

• Capture the potential failure modes and failure sequence of the truss under combined gravity

load and thermal load;

• Develop an understanding of the relative importance of different structural features and

failure modes; and

• Develop a simplified model that replicates the expected failure and the limit loads of the truss

to be used for analysis of the full floor subsystem model.
Interim Report on Subsystem Structural Analysis of the WTC Towers

Beam element

Rigid, beam

Rigid beam

Beam element

Control element No. 1: Capture

walk-off support

B,[(2,3,UZ);(2,l,UY);(K,Ao)]

Control element No. 2: Capture seat vertical shear capacity

B:[(l,3,UZ);(4,2,UZ);(K,Ao)]

Control element No. 3:

Capture loss of horizontal

resistance if seat fails vertically

B3[(2,3,UY);(4,2,UZ);(K,Ao)]

Control element No. 4:

Capture bolt shear capacity

B4[(1.3,UY);(2,5,UY);(K,Ao)]

Beam element

(Out-of-plan)

Beam element No. 1: Make seat vertical shear capacity temoerature-deoendent

Beam element No. 2: Make bolt shear capacity

temperature-dependent

Constraint equations

Coupling displacement

DOF of node 1 and 6

Beam element

(Out-of-plan)

Y

Coordinate system

Figure K-31. Simplified model of interior seat.
Appendix K

(xlO**l)

Truss walks off

the support

0 1 2 3 4 5

.5 1.5 2.5 3.5 4.5 Truss travel distance, in.

Figure K-32. Results of simplified seat model capturing failure

from truss walking off interior seat.

(xlO**l)

0 1 2 3 4 5

.5 1.5 2.5 3.5 4.5 Truss travel distance, in.

Figure K-33. Results of simplified seat model capturing failure

from exceeding the interior seat vertical shear capacity.

!

I
Interim Report on Subsystem Structural Analysis of the WTC Towers

Exterior seat model same as the

interior seat

Beam element: Make gusset plate tensile strength

temperature-dependent

Fix rotational DOF

Constraint equations

Coupling displacement DOF of node 1 and 3

Breali element:

Capture failure of gusset plate under tensile force

B|[(l,2,UY);(4,2,UY);(K,Ao)]

Y

<-

AZ

Coordinate system

Figure K-34. Simplified model of exterior seat.

K-45
Appendix K

fl-, - Node on slab

Five beam elements are used to make the knuckle

capacity

temperature-dependent

Point-to-point contact element to transfer

vertical compressive force between node 1 and 2

Constraint equations

Coupling displacement

DOF of node 1 and 7

Break element No. 10: Capture loss of horizontal

resistance in the Y direction if knuckle fails vertically

B,o[(2,6,UY);(12,2,UZ);(K,Ao)]

Break element No. 1-4: Capture loss of vertical resistance if knuckle fails horizontally

B|[(2,6,UZ);(8,2,UX);(K,Ao)]

B2[(6,5,UZ);(2,9,UX);(K,Ao)]

B,[(5,4,UZ);(10,2,UY);(K,Ao)]

B4[(4 3,yZ);(2 ll,UY);(K,Ao)]

Break element No. 5: Capture knuckle tensile failure

B5[(3,l,UZ);(12,2,UZ);(K,Ao)]

The purpose of break

elements No. 11-15 is similar to break elements No. 6-10

Break element No. 8-9: Capture loss of horizontal

resistance in the Y direction if knuckle fails horizontally in the X

direction

B«[(5,4,UY);(8,2,UX);(K,Ao)]

Bor(6.5.UY):(2.9.UX):fK.Anl1

Break element No. 6: Capture knuckle horizontal compression failure in the Y direction

B6[(3,l,UY);(10,2,UY);(K,Ao)]

Break element No. 7: Capture knuckle horizontal tensile

failure in the Y direction

Bv[(4,3,UY);{2,ll,UY);(K,Ao)]

Figure K-35. Simplified model of knuckle.

K-46
Interim Report on Subsystem Structural Analysis of the WTC Towers

Failure Modes

The model can capture the following failure modes:

Softening and Sagging of Truss—-The top and bottom chords and diagonals of the truss are exposed to the hot gas layer below the floor slab. As described in Section K.5, the steel in the truss exhibits stiffness degradation, yield strength reduction, plastic softening, and creep at high temperatures. A truss with

softened chords sags. The heat may also reduce the stiffness and strength of the concrete slab, especially

its bottom layer where temperature is the highest, and around the knuckle where concrete temperature

rises by conduction through the steel.

In addition to direct thermal effects, sagging and weakening of the truss can be caused by the following

failure modes:

• Buckling or failure of web diagonal members, which reduces the truss action and causes the

truss to act as a catenary;

• Buckling or failure of the top and bottom chord members;

• Knuckle failure and loss of composite action of the concrete slab and the steel truss; or

• Weld failure between the diagonal and the chord.

Loss ofSupport of Truss—The truss can fail by loss of support due to seat failure. Loss of support at either the exterior or interior seat can be caused by the extreme sagging and catenary action of the truss due to plastic deformation and buckling of truss members.

As discussed under Boundary Conditions later in this section, the bottom chord of the truss is restrained

in the lateral direction at the bridging tmss locations. Although the out-of-plane deformation of the bottom chord due to thermal expansion of bridging trusses will result in a reduction in the vertical load

capacity of a primary truss, the truss model studied here cannot capture this phenomenon. The interaction

between the bridging trusses and the primary trusses is intended to be captured in the full floor model.

Model Description

Figure K-36 shows the truss model. A typical long-span truss designated C32T1 (SHCR 1973:WTC

Drawing Book 7, Sheet ABl-2) is modeled to study its response to failure when subjected to dead and

live loads and thermal loads. The model includes the following:

• One truss of the pair of trusses at column line 143 of floor 96 in WTC 1

;

• Two exterior columns (columns 143 and 144) with half the area and bending properties, and a length of 24 ft (12 ft above and below the floor level);

• The portion of the spandrel between the two exterior columns;

K-47
Appendix K

288 in.

span=713 in.

(a) Entire model

knuckle

top chord

<——r-r-nr —

r— '1——

4.35 in.

slab

top chord

bearing angle

web

bottom chord

29 in.

bottom chord

(b) Cross section (c) Top and bottom chords and web diagonals

Figure K-36. Truss model.

• The portion of the slab (40 in. wide) between the two exterior columns;

• One strap anchor that is attached to the truss top chord, concrete slab and the adjacent exterior

column (column 144); and

• Exterior and interior seats, and the gusset plate at the exterior end.

A typical slab section consists of 4 in. thick lightweight concrete on 22 gauge metal deck with flutes 6.8 in. on centers. Two layers of welded wire fabric were provided in the slab. The reinforcement ratios are 0.21 and 0.735 in the directions along and transverse to the truss, respectively. A flute is 2 in. wide at the top, 1.25 in. wide at the bottom, and 1.47 in. high. An equivalent thickness of 4.35 in. is used as the

slab thickness in the truss model. By using the equivalent thickness, the bending stiffness in the direction transverse to the truss is about 1 5 percent higher than the actual stiffness. However, since the bending in the transverse direction in this truss model is insignificant, the slab is modeled as an isotropic plate. The

metal deck and the welded wire fabric are not included in the truss model.

The top and bottom chords and the diagonals of the truss are modeled by 3-D quadratic finite strain beam

(BEAM! 89) elements with temperature-dependent elastic, plastic, and creep material properties. The top
Interim Report on Subsystem Structural Analysis of the WTC Towers

chord consists of double angles of 1 1/2 x 2 x 0.25 (long legs horizontal), while the bottom chord consists

of double angles of 3 x 2 x 0.37 (long legs horizontal). Web members are round bars of either 1 .09 in. or

1.14 in. diameter. A typical diagonal member has a 1 .09 in. diameter. Top and bottom chords are divided

into four elements between panel points, and a diagonal is also divided into four elements between top

and bottom chords. The concrete slab is modeled with 4-node finite strain shell (SHELL181) elements.

The nodes of the concrete slab are located atthe neutral plane of the concrete slab with an offset relative

to the nodes of the top chords. The cast iron model (Hjelm model) can be used with the SHELL181

elements that allow different "yield" in tension and compression. A low "yield stress in tension" is used

to simulate cracking.

At knuckle locations, the top chord elements and the elements representing the concrete slab are connected by control elements (COMB1N37) with capacities determined from the detailed knuckle

analysis. By including point-to-point contact (CONTA178) elements, compression can always be

transferred even after knuckles fail. Studs on the strap between the top chord and column 144 are also modeled by COMBIN37 elements that connect the strap to the slab and have temperature-dependent

capacities. The slab and the strap are tied by the COMB1N37 elements horizontally while their vertical displacements are coupled. The exterior and core seats are modeled by COMBIN37 elements that have

temperature-dependent capacities determined from the seat analysis. A stud on the spandrel is also modeled by a COMB1N37 element, which ties the spandrel with the slab and has temperature-dependent

capacities. Because only one 5/8 in. stud was provided over 80 in. between the slab and the spandrel, a half stud capacity is considered for the stud in the model. In addition, the first stud on the strap is located near this stud on the spandrel. Therefore, COMBIN37 elements between the slab and the spandrel have a capacity of a combination of these studs, including a group effect. Damping unit connecting the truss bottom chord to the spandrel plate is assumed to have little effect on the behavior of the floor truss under

sustained loading; therefore, it was not included in the model.

Three-D elastic beam (BEAM44) elements model the exterior columns. SHELL63 elastic shell elements

model the spandrel.

Boundary Conditions

Boundary conditions on the truss model are shown in Fig. K-37.

The entire top chord of the truss is supported in the x direction. The bottom chord is supported in the x

direction at four bridging truss locations. Two edges of the concrete slab are restrained against rotations

about the y and z axes, but can move in the x direction.

The interior truss seat is fixed in all directions. The exterior seat is fixed to the spandrel. The truss is pinned at both exterior and interior truss seats.

The exterior end of the slab is tied to the spandrel by only COMBIN37 elements representing studs. The

interior end of the slab is fixed in the z direction and in rotation about the z direction. In the y direction at the interior end of slab, break elements that have temperature-dependent tensile capacities are implemented as show in Fig. K-38. Therefore, the interior slab end is fixed in the y direction until the

tensile force exceeds the capacity that is calculated based on the amount of steel reinforcement (#3@10"

top and #4@\2" bottom).

K-49
Appendix K

Fixed UZ, UY,

and ROTY at

Figure K-37. Boundary conditions.

Figure K-38. Break elements at the interior end of slab.

In the analysis with increasing gravity load, a model different from the current model is used, where

boundary conditions of the slab are slightly different from the current model described above. In this model, the exterior end of the slab is tied to the spandrel, and the core end of the slab is fixed in the v and

z directions and in rotations about the x and z directions. Also, COMBIN37 elements for seats are not included in this model.

Loading

The loading on the truss model consists of gravity dead and live loads and temperature time-histories for

all steel members, including the truss seats. The gravity loads include weight of the structure, 8 psf

superimposed dead load (including nonstructural dead loads due to architectural items and fixed service

equipment), and 13.75 psf of live load equal to 25 percent of design live load of 55 psf. The thermal load

is a linear temperature gradient through the slab fi-om 300 °C at the top surface of the concrete slab to

K-50
Interim Report on Subsystem Structural Analysis of the WTC Towers

700 °C at the bottom surface of the slab. The temperature is ramped from 20 °C to 700 °C in steel members; from 20 °C to 700 °C at the bottom surface of the slab and from 20 °C to 300 °C at the top

surface of the slab at 1,800 s; thereafter, the temperatures do not change for another 1,800 s. Temperature

is not applied to the columns.

In order to determine the effect of debris load on the truss behavior, a parametric study will be performed.

Material Properties

Table K-12 shows material assignments for different structural components in the truss model.

Table K-12. IVIaterial assignments in truss model.

Structural Component

Specified Yield

Strength Material ID Top chord 50 ksi 21 Bottom chord 50ksi 21

1.09 in. diameter web 36 ksi 20

1 . 14 in. diameter web 50 ksi 21

Strap 36 ksi 1 Column 143 65 ksi 15 Column 144 65 ksi 15

Spandrel 42 ksi 1 1

Lightweight concrete slab 3,000 psi (f c) 83

Columns 143 and 144 and the spandrel, use only elastic properties. In the current model, the concrete

slab also remains elastic.

Resistance Welds

Table K-13 shows the resistance weld strength between a chord (double angles) and a diagonal based on

the test data found at Laclede Steel. Weld strength shown in Table K-13 is the sum of the capacities of

two resistance welds. Figure K-39 compares resistance weld strength between top or bottom chord and a diagonal with yield strength of a diagonal at elevated temperatures. As can be seen in Fig. K-39 (a), a

typical diagonal (1.09 in. diameter) will yield before the resistance weld fails. For 1.14 in. diameter

diagonal, the resistance weld strength cannot yield the bar at temperatures below 550 °C, as can be seen in Fig. K-39 (b). However, shop drawings show additional arc welds between the chord and 1.14 in. diameter bar at most locations.

Table K-13. Resistance weld strength.

Chord Diagonal Size (in.) Average Strength (kip)

Top chord 1.09 36.9 Top chord 1.14 37.7 Bottom chord 1.09 41.0 Bottom chord 1.14 40.5

K-51
Appendix K

50

40

\**.

* \'

V V

* \

60

40

200 400 Temperature (°C) Top Chord

Bottom Chord

Yield Strength

600 800

* — \

••-.^ \

. ... h- -

V

0 200 400 Temperature (°C) Top Chord

Bottom Chord

Yield Strength

800

(a) 1.09 in. diameter bar (b) 1.14 in. diameter bar

Figure K-39. Comparison of resistance weld strength and yield strength of web member

at elevated temperatures.

Current Status

The truss model can capture the following: / ,

• Temperature-dependent elastic material properties for both steel and concrete;

• Temperature-dependent steel plasticity;

• Buckling of truss members;

• Failure of knuckle—loss of composite action;

• Failure of studs on the strap;

• Failure of stud between the spandrel and the concrete slab; and

• Failure of the exterior and interior truss seats.

The following features are being added to the truss model:

• Crushing and cracking of concrete;

• Creep strain in steel at elevated temperatures; and

• Failure of welds (Calculations show section yielding can occur prior to weld failure in nearly

all cases.).

K-52
Interim Report on Subsystem Structural Analysis of the WTC Towers

Model Verification

The maximum vertical displacement is checked against the single truss model extracted from the ANSYS

full floor model that was converted from the SAP full floor model. The difference in the vertical displacement is only 3.5 percent.

FEA Results

Gravity Loading—The maximum calculated vertical deflection is 1 . 1 in. downward. The maximum

calculated horizontal column deflection is 0.022 in. inward. The maximum forces in top chord, bottom

chord, and diagonal are 13,357 lb, 39,514 lb, and 7,647 lb, respectively.

Gravity Plus Thermal Loading—The analysis is carried out dynamically with 5 percent Rayleigh

damping. To shorten the run time, the total time period is set to 1.0 s for the temperature ramp. The

analysis proceeded to a temperature of r=663 °C. Figure K-40 shows horizontal and vertical displacement results. A positive horizontal displacement indicates that the exterior columns are pushed

out, and a negative vertical displacement indicates that the truss is deflected downward. At 340 °C, the horizontal displacement at the exterior column starts to decrease. At 560 °C, the exterior columns are pulled in, and the truss becomes catenary from that point on.

800 Temperature (°C) Temperature (°C) (a) Horizontal displacement at column 143 (b) Vertical displacement at midspan

Figure K-40. Displacement versus temperature.

K-53
Appendix K

Figure shows axial forces in the truss members. In the figure, Py is the axial force at yield and

equals the product of the net area of the member and the yield strength which varies with temperature. Pc

is the compressive strength per AISC formula (AISC 2003) for the top chord with fixed ends in Fig. K-41 (a) and for 1.09 in. diameter diagonal bar with pinned ends in Fig. K^l (c).

20000

0

5 -20000 o

u

O -40000

^ -60000

-100000

—B-^EL#8 —1— EL#13 —o— EL#61 —«— EL#86 —»— EL#92 —^— EL#100

Py

100000

80000

£ 60000

«>u

O 40000

u.

"(5

]| 20000

0

-20000

-EL#168 (mid) -EL#120 (nearext) -EL#199(near int) Py

100 200 300 400 500 Temperature (°C)

(a) Top chord

5000

0

700 100 200 300 400 500 Temperature (°C)

(b) Bottom chord

600 700

£ -5000

O -10000

>< -15000

-20000

-25000

-—o— EL#222 EL#236 - —°— EL#357 EL#371 !^»— EL#385 —»— EL#399 —9— EL#413(1.14in) Py

1 PcrforL=32 in

100 200 300 400 500 Temperature (°C)600 700

(c) Compression diagonal

Figure K-41. Axial force In truss members versus temperature.

Figure K-42 (a) shows the top chords yielding beyond 300 °C. This is due to a significant difference of

coefficients of thermal expansion (CTE) between concrete and steel. At 500 °C, the CTE of steel is twice

that of lightweight concrete. Bottom chords are still in the elastic range at the end of analysis. Some

diagonals are bent significantly in the plane of the truss by high axial force and end moments (see

Fig. K-42 for the deformed shape at the interior end). This diagonal buckling starts at approximately

340 °C.

K-54
Interim Report on Subsystem Structural Analysis of the WTC Towers

-23549 -13246 -2944 7359 17661 -18398 -8095 2207 12510 22812

Figure K-42. Axial stress contour in the truss members at 663 °C

(displacement magnification factor = 1.0).

Figure shows knuckle forces in the y direction (longitudinal truss direction) and the z direction

(vertical direction). The capacity of a knuckle in the y direction is assumed to be 30,000 lb, and in the

z direction 15,000 lb in tension. Knuckles 14 and 15 fail due to horizontal shear around 400 °C.

Knuckle 1 also fails due to the horizontal shear around 650 °C.

30000

-30000

10 11 12 13 14 15

Knuckle 1

Knuckle 2

Knuckle 3

Knuckle 13

Knuckle 14

Knuckle 15

200 400 Temperature (°C)

600

Knuckle 1 failed

800 200 400 Temperature (°C)

600 800

(a) Horizontal Force (b) Vertical Force

Figure K-43. Force in the knuckles versus temperature.

Figure K-^4 (a) and (b) show horizontal and vertical reaction forces at seats, respectively. At 510^0, the

interior seat bolt shears off At 650 T, the truss walks off the interior seat. At 660^0, the gusset plate at the exterior end fails in tension.

K-55
Appendix K

Temperature (°C) Temperature (°C)

(a) Horizontal reaction force . (b) Vertical reaction force

Figure K-44. Reaction forces at seats.

Additional Debris Load

The capacity of the truss model against additional debris load is determined by increasing the gravity

loading at room temperature. The analysis is performed with the previous model, where boundary

conditions of the slab are as described in the section "Bounding Conditions." Let us define load factor as the ratio of the gravity load plus debris weight to the gravity load, where gravity load includes self weight,

superimposed dead load, and 25 percent of the reduced live load. The analysis was terminated at a load

factor of 3.4. Figure K^5 (a) shows midspan vertical displacement versus load factor. At 2.4 times the

gravity loading, 1 1 knuckles from the core end fail in the truss direction. At 2.8 times the gravity loading,

the fourth knuckle from the exterior end fails. Figure K-A5 (b) shows the sum of horizontal reaction

forces measured at the exterior columns. Note that seat capacities are not modeled in this analysis.

1. 2 3 4 5 6 7 8 9 10 11 12 13 14 15

Displacement (in) Horizontal Reaction Force (lb)

(a) Load factor versus midspan vertical (b) Load factor versus horizontal reaction

displacement ^-z force

Figure K-45. Finite element analysis results from increasing gravity.

K-56
Interim Report on Subsystem Structural Analysis of the WTC Towers

Summary and Discussion

The truss behavior under the gravity plus thermal loading, where the temperature is ramped up to 663 °C

can be summarized as follows:

• Top chords yield above 300 °C due to the difference in CTEs of steel and lightweight

concrete.

• Compression diagonals start to buckle in the plane of the truss due to a high axial force and

end moments at 340 "C.

• At 400 °C, knuckles start to fail.

• The interior seat bolt shears off at 5 1 0 °C.

• The truss walks off the interior seat at 650 °C, followed by fracture of the gusset plate at the

exterior end at 663 °C.

The results for the additional debris weight show that the knuckles start failing when the load factor is 2.4. Most knuckles fail before load factor reaches 3.0. After the knuckle failure, the truss loses composite

action between the truss and the concrete slab, and the vertical displacement increases significantly. As a

result, horizontal reaction force increases.

Models of the truss including knuckles with temperature-dependent capacities, diagonal weld failure, and

concrete cracking and crushing are under study.

Simplified Model

To be used in the full floor subsystem model, the truss model will be simplified based on the results from

the truss model analysis. Characteristics of the simplified truss model are listed in the following:

• The geometry of the truss will be preserved.

• Pin-ended Link elements will be used for truss members.

• User-defined elements will be used to model failure modes of knuckles, seats, and diagonal

members. They will be implemented at the ends of link elements.

• Slab softening or cracking will be incorporated into the model.

K.8 EXTERIOR WALL SUBSYSTEM

The exterior wall subsystem represents the impact zone and includes nine prefabricated wall panels, three panels high by three panels wide.

The exterior wall subsystem model includes nine columns, extending vertically from the column splice

located below floor 91 to the column splice above floor 99, and nine spandrels, extending horizontally

K-57
Appendix K

from the spandrel splice located at mid-span between columns 149 and 150 to the spandrel splice at mid-span between columns 158 and 159, of the WTC 1 exterior wall.

Figure shows the subsystem pictorially. Tables K-14 through K-16 give the properties of the

column component plates, the spandrels, and the column and spandrel splices. Figure K—47 shows

pictorially the spandrel plate thickness, nominal yield strengths, and spandrel splice types. Figure K-48

shows the column plates notation used.

The odd numbered columns support floor trusses. Pairs of strap anchors extend diagonally from the top

chord of truss pairs to the even numbered columns. The trusses and the straps partially brace the columns

both in-plane and out-of-plane of the exterior wall.

Figure K-46. Exterior wall subsystem structure.

K-58
Interim Report on Subsystem Structural Analysis of the WTC Towers

Column - 150 151 152 153 154 155 156 157 158

Floor - 99

1

98:

97 r

961.

95i94 r

93t92 L

91r

XXX-XX - Indicates type and nominal yield strength of column

XXX - Indicates type of column splice

a) Columns and Column Splices

Column - 150 151 152 153 154 155 156 157 158

I

I

Floor - 991^^

98 im.

97 r

95l^

94r

93 f

92 C

91^

-45

-50

-45

-42

-45

-42

-45

-42

m.

102

102

-45

-46

-50

-45 102

-46

-50

-45

-46

-50

-50

-55

-60

-65

-65

-70

-70

XXX-XX - Indicates nominal yield strength of spandrel XXX - Indicates type of spandrel splice Spandrel Thickness 3/8 in.

b) Spandrels and Spandrel Splices

Figure K-47. Column and floor number materials and splice types.

K-59
Appendix K

ECRL

Plate 1

14"

Plate 2 n

Plate 1

Plate 3

13.5"

Spandrel

Figure K-48. Spandrels and spandrel splices.

Column Plate 1 Plate 2 Plate 3 Col. Type

Type 1 X t 1 X t 1 X t ID

120 13.5 X 0.25 13.5 X 0.25 15.75 X 0.25 0

121 13.5 X 0.3125 13.375 X 0.25 15.75 X 0.25 1

1 22 13.5 X 0.375 13.25 X 0.25 15.75 X 0.25 2

123 13.5 X 0.4375 13.125 X 0.25 15.75 X 0.25 3 124 13.5 X 0.5 13 X 0.25 15.75 X 0.25 4

125 13.5 X 0.5625 12.875 X 0.25 15.75 X 0.25 5 Note: All spandrels in wall model are 52 in. deep 3/8 in. thick.

Table K-15. Spandrel splice details.

Spandrel

Splice

Type

Number of

Bolts/Row

Total

Number of

Rows

Bolt Spacing Gage

Overall Splice

Plate Dimensions

Bolt to Centerline of

Splice

Gap B/W

Spandrels

Spandrel

Splice ID

101 6 2 5@9 49 X 6.75 X.25 1.875 0.75 101

102 8 2 3,6,3@

9,6,3

49 X 6.75 x.25 1.875 0.75 102

111 6 4 5@9 3 49 X 12.75 x.25 1.875 0.75 111

112 8 4 3,6,3@

9,6,3

3 49 X 12.75 x.25 1.875 0.75 112

a. All spandrel splices use 7/8 in. A325 bolts; spandrel plate yield strength is 36 ksi. b. Holes in spandrel are 1/4 in. larger than bolts; holes in plates are bolt + 1/16 in. or option to match spandrel holes.

Table K-16. Column splice details

Column

Splice Type

Butt Plate Thickness

Number

of Bolts

Bolt

Diameter Gage

Bolt SpacingColumn

Splice ID

411 1.375 4 0.875 3.5 6 411

421 1.625 4 0.875 3.5 6 421

431 1.875 4 1 3.5 6 431

a. Butt plates have specified yield strength of 50 ksi.

b. Bolts are A325.

K-60
Interim Report on Subsystem Structural Analysis of the WTC Towers

K.8.1 Description of Exterior Wall Subsystem Model

Figure K-49 shows the model in elevation. BEAM189 elements model the columns. SHELL181 plate

elements model the spandrels. Figure K-50 shows the number of elements used to model columns and

spandrels. MFC 184 rigid elements connect the center of gravity of a column to the mid-plane of a spandrel at each shell element. Figure K-5 1, shows this use of the MFC 1 84. MFC 1 84 rigid elements also model the spandrel connections. A simplified model, consisting of two BEAM189 elements for each of

the four bolts, four pairs of CONTA178 contact elements at the faying (contact) surfaces, and MFC 184

rigid elements connecting the tops of the bolts to the CONTA178 contact elements, model the column

splice. COMBIN37 elements model the firacture of the column splice bolts.

Z

Floor 99

Y

X

Floor 91

UY,UZ supports, typ

-» UY supports at top, typ.

UY support at truss connections

(gusset plates) - each floor, each

column

ROTY supports at spandrel edges, typ.

UX,UY,UZ support at middle column

Figure K-49. Exterior wall subsystem model, viewed from inside of WTC 1.

K-61
Appendix K

10 elements

through depth

4 elements each side of splice

SHELL181

elements for spandrels

BEAM189

elements for columns

2 elements at columns

4 elements

between

columns, typ.

17.33 in. t (gusset plate)

25.5 in. (center

of stand-off)

Column

splices

1 element at spandrel

splice

Plate 1

-

t varies

Figure K-50. Portion of exterior wall subsystem model showing

number of elements used.

Plate 2

t=0.25 in.

-Plate 1

t varies

Plate 2

t=0.25 in. L

Plate 1

t varies

Plate 1

t varies

7

Plate 4

t=0.005 in. — MPC184 rigid beams

attach spandrel to column e.g. to create

composite action

At spandrels between column and

spandrel.

"T

Plate 3

t=0.25 in.

Away from spandrels

Figure K-51. Schematic representation of columns used in the exterior wall

subsystem model.

The capabilities of the BEAMI89 and SHELL181 elements include large deflections, plastic deformation,

and creep at elevated temperatures. Materials are assigned as described in Section K.5.

The loads on the model include the following:

• Self weight;

• Dead load of floor trusses;

K-62
Interim Report on Subsystem Structural Analysis of the WTC Towers

• 25 percent of floor live load;

• Column splice bolt preload; and

• Temperatures of fire scenarios.

A concentrated vertical load and an out-of-th'e-wall-plane moment due to the dead and live load of the

structure above floor 99 load the top of each column. A concentrated vertical load and an

out-of-the-wall-plane couple due to the dead and live load of the floor truss load each odd numbered

column at the truss seats. Mean temperature at the center of gravity of the column and a linear gradient in each of two directions through the section of the column strain the BEAM189 elements at each node.

Temperatures at the nodes strain the SHELL181 elements. Loads and/or deflections at the truss seats model the outward motion or the catemary action of the floor truss due to fire scenarios. The

7/8 in. diameter column splice bolts are preloaded with 36.05 kip (AISC 1964).

Simple supports out of the plane of the wall restrain the tops and the bottoms of all columns in the model.

In addition, supports horizontally in the plane of the wall restrain the top and the bottom of central column 154. Simple supports in the vertical direction restrain the bottoms of all columns in the model.

Symmetry conditions are imposed on the spandrels at the extremities of the model, except that the spandrels are free to expand in the plane of the wall. In the plane and out of the plane of the wall

restraints brace the column at floor truss seats and diagonal straps.

The model captures the following failure modes:

• Column collapse due to large lateral deformations;

• Column buckling due to loss of bracing at floor truss seats and diagonal straps;

• Failure of column splice bolts; and

• Failure of spandrel splice bolts.

The model does not capmre the local buckling of the column plates and the formation of plastic hinges

due to the interaction of local plate buckling and high stresses in the column from axial load and bending

moments. Section K.8.4 below includes the justification for excluding this structural behavior from the wall subsystem model.

K.8.2 Validation of the Exterior Wall Subsystem Model

The behavior of models of the following components of the exterior wall subsystem validate the exterior

wall subsystem model:

• Model of a one-story-high exterior column.

• Model of a nine-story-high exterior column.

K-63
Appendix K

• Detailed and simplified models of the column splice.

• SAP2000 and ANSYS models of a prefabricated wall panel.

K.8.3 Model of One-Story High Exterior Column

Figure K-52 shows the model of a one-story-high exterior column. The model includes a one-story-high

portion of column 1 5 1 extending from floor 95 to floor 96 and portions of spandrels at floor 95 and

floor 96. The model also represents column 151 from floor 96 to floor 97 since the dimensions, plate

thicknesses and material properties are identical to those of column 151 from floor 95 to floor 96. SHELL181 plate elements model the plates of the column and the spandrels. CERJG rigid elements

connect the center of gravity of the column to its component plates and the spandrel at both the top and

bottom of the model. The column is simply supported in three directions at the bottom and simply

supported in the horizontal direction at the top. Increments of axial displacement applied at the top load

the model.

ANSYS

MAR 2 9 2004

13: 49: 28

Figure K-52. One-story exterior column model.

Figure K-53 shows the variation of axial load with enforced axial displacement and resulting lateral deflection at room temperature and 700 °C. This figure also shows the hand calculated column load

levels at room temperature and 700 °C for:

• Local buckling of Plate 2 and Plate 3;

• Unifonn yielding of the column; and

• Axial load due to dead and live load at floor 96 in the exterior wall subsystem model.

K-64
Interim Report on Subsystem Structural Analysis of the WTC Towers

Figure K-53. Load-deflection of column at room temperature and 700 °C.

Figure K-54 shows the local bucking deformation of Plate 2 and Plate 3 at the maximum load level. Figure K-55 shows a plastic hinge at mid-height of the column for an axial displacement of 2 in. Figure K-56 shows the presence of local buckles in Plate 2 and Plate 3 at the maximum load.

Figure K-53 shows that at room temperature Plate 2 and Plate 3 buckle locally at a load that is less than

the maximum column load, but that at 700 °C the column yields before it buckles locally. This figure also shows that the expected column demand load of 1 75 kip is substantially lower than the local buckling

load at room temperature and the column yield load at 700 °C. For these results, the axial displacement was applied along the center of gravity of the column cross section away from the spandrel. If axial

displacement is applied at center of gravity of the column cross section at the spandrel, there will be

additional bending moment in the column section away from the spandrel. The presence of moments

reduces the axial load capacity of the column. The resulting load-deflection diagram is also shown in Fig. K-53.

K.8.4 Model of Nine-Story High Exterior Model

Figure K-57 shows the nine-story-high exterior column model. The model includes column 151

extending from near mid-height between floor 91 and floor 92 to mid-height between floor 100 and

floor 101, spandrels at floors 92 through 100, and column splices located at the mid-height between

floors 94 and 95 and floors 97 and 98. SHELL181 plate elements model the plates of the column, the spandrels, the butt plates at the column splice, and the stiffeners. BEAM189 elements model the column

splice bohs. CONTA174 and TARGE170 elements model the faying surfaces of the column splice.

MPC184 rigid elements connect the tops of the bolts to the butt plates. At the bottom the column is restrained from displacement and rotation in all three directions. At the top the column is restrained from

translating in the horizontal directions and from twisting.

K-65
Appendix K

DISPLACEMENT

STEP=3

SUB =34 TIME=1. 346

DMX =.36137

ANSYS

MAR 2 2004

14:27: 10

Figure K-54. Local buckling of column at room temperature.

1 DISPLACEMENT

STEP=4

SUB =200

TIME=3 DMX =9. 169

ANSYS

MAR 2 2004 14:31:48

DISPL. UP TO 2in

Figure K-55. Plastic hinge in column at room temperature.

K-66
Interim Report on Subsystem Structural Analysis of the WTC Towers

ANSYS

MAR 29 200'! 13: 45: 08

Figure K-56. Deformed shape of column at maximum axial load at 700 °C.

Figure K-57. Nine-story column model.

K-67
Appendix K

The capabilities of the BEAM189 and SHELL181 elements include large deflections and plastic

deformation. For these elements, gives the material property identification numbers, which in turn are described in Chapter 4 above.

The loads on the model include the following:

• Self weight;

• Dead load of floor trusses;

• 25 percent of floor live load;

• Column splice bolt preload; and

• Temperature of Fire Scenario G.

In Fire Scenario G, the fire starts on floors 95, 96, and 97 and spreads to floors 93 through 98. Gas

temperature reaches 1,100 °C. Convection cools the outside face of the column. Radiation heats the

other three faces. The inside face of the column is not fireproofed. Temperatures are provided at 200 s intervals up to 5,000 s. Figure K-58 shows the variation of the maximum temperature anywhere in the

column with time and the yield stress at the point of maximum temperature. The temperature reaches a maximum of 706 °C at 5,000 s.

70000 -

" 40000 -

10000

2400 2800 Time (sec)

Figure K-58. Variation of maximum temperature and

corresponding yield stress with time, fire scenario G.

To account for the dead and live load of the structure above floor 100 and of the floors that connect to the

column, concentrated vertical loads and bending moments about a horizontal axis in the plane of the wall

are applied to the top of the column and at all truss seats. Furthermore, the 7/8 in. diameter column splice bolts are preloaded to 36.05 kip (AISC 1964).

K-68
Interim Report on Subsystem Structural Analysis of the WTC Towers

Figure K-59 shows the variation of maximum tensile and maximum compressive stresses with time and

the corresponding yield stress. Figure K-60 shows the deformed shape of the column at 400 s when the compression stress in the column is a maximum. Figure K-60 also shows the deformed shape of the column at 3,200 s when the tensile stress in the column is a maximum. Figure K-60 also shows the deformed shape of the column at 5,000 s when the temperature in the column is a maximum.

Tension

£0

Compression

52 DO

Time (Sec)

Figure K-59. Maximum compressive and tensile axial stress and

corresponding yield stress with time, fire scenario G.

Figure K-59 shows that the tensile and compressive stresses exceed the yield stress for most times during

the duration of the fire.

Figure K-60 shows that for Fire Scenario G, an extreme scenario that assumes no fireproofmg on the

inside face of the column, plastic hinges do not form in the column. This justifies the exclusion of local buckling of the column plates from the wall subsystem model.

K.8.5 Models of the Column Splice

The plate model of the column splice, shown in Fig. K-6 1 includes a 92 in. tall section of the nine-story

column model centered on the column splice located below floor 98.

Figure K-62 shows a simplified model of the column splice. The simplified model consists of two

BEAM 189 elements for each of the four bolts, four pairs of CONTA178 contact elements at the faying

surfaces, and MFC 184 rigid elements connecting the ends of the bohs to the CONTA178 contact elements. A BEAM 189 element extends from each side of the splice to match the length of the plate

model. Figure K-63 shows the details that model the faying surfaces of the splice.

K-69
Appendix K

ANSYSI

MAR 3 0 2004

09: 43: 21

a)Tiine 400$ b)Time SZOOs c)Time 5000s

Figure K-60. Deformed shape of column at 400 s, 3,200 s, and 5,000 s (floors 95-97).

ELEMENTS

92

ANSYS

FEB 19 2004

08: 57: 19

Figure K-61. Plate model of column splice, floors 97-98.

K-70
Interim Report on Subsystem Structural Analysis of the WTC Towers

ELEMENTS ANSYSI

MAR 2 9 20 0 4 13: 51: 07

Figure K-62. Simplified model of column splice.

Figure K-63. Column splice details, plate model and simplified model.

Both models are subjected to the following loads:

• Axial tension;

• Shear transverse to the plane of the wall;

• Moment out of plane of the wall;

• Moment in plane of the wall; and

• Torsion.

K-71
Appendix K

Figure K-64 shows the variation of axial displacement with axial force load. Figure K-65 shows the variation of transverse displacement with transverse shear fore. These figures show excellent agreement

of the simplified model with the plate model.

80000

0.005 -0,015 -0.035 -0.055 -0.075 -0.095

Vertical Displacement (in)

-0.115 -0.135 -0.155 -0.175

Figure K-64. Variation of axial displacement with axial load.

JUUUUU

Lateral Displacement (in)

Figure K-65. Variation of lateral displacement with shear load.

K-72
Interim Report on Subsystem Structural Analysis of the WTC Towers

Figure K-66 shows the rotation variation with out-of-plane of the wall moment. Figure K-67 shows the

rotation variation with in-plane-of-the-wall moment. Figure K-68 shows the twist variation with torque.

Figures K-66, K-67, and Fig. K-68 show large differences between the results of the simplified and plate

models of the column splice. These differences are due to the fixed locations of pivot points in the simplified model, provided by pairs of CONTA178 point-to-point contact elements, about which the faying surfaces rotate. The CONTA174 and TARGE170 surface contact elements for the faying surfaces

in the plate model permit the location of the pivot point to adjust to the demand of the applied moment.

Adjusting the location of the point-to-point contact elements can minimize these differences, but they

cannot be eliminated. In the exterior wall subsystem model, the locations of the single point contact elements in the column splices will be adjusted and the sensitivity of the response of the model results to these locations computed.

1800000 1 ! ] , , 1 1

1600000

Rotation (radians) Figure K-66. Variation of rotation with moment, out of wall plane.

Rotation (radians) Figure K-67. Variation of rotation with moment, in plane of wall.

K-73
Appendix K

2000000

1800000

1600000

0.00 0.05 0.10 0.15 0.20 0.25 0.30 Twist Angle (radians) Figure K-68. Variation of twist angle with torque.

0.35 0.40

K.8.6 Prefabricated Panel Model

Description of Model

Figure K-69 shows the SAP2000 model of a typical prefabricated panel at floors 79 to 82. The model is modified as follows:

• Eliminated self-weight from loading conditions.

• Provided stiff members at the tops of the columns and replaced the four concentrated loads

with a single concentrated load.

• Added out-of-plane of the wall supports (UY) at top of columns for out-of-plane loading.

Figure K-70 shows the ANSYS model for matching the behavior of the SAP2000 exterior wall

subsystem model. In the ANSYS version of the panel model BEAIVI189 elements model the columns and

MPC184 rigid elements attach the spandrels to the columns.

K-74
Interim Report on Subsystem Structural Analysis of the WTC Towers

Figure K-69. SAP2000 model of prefabricated panel.

46 in.

52 in.

92 in.

52 in.

92 in.

52 in.

46 in.

-Rigid truss members to transfer horizontal forces

20 in 20 in.

t = 0.4375 in., typ

10 elements between spandrels

9 elements deep, 4 elements

between columns

- t = 0,4370 in. at columns,

2 elements wide

Figure K-70. ANSYS model of prefabricated panel showing

geometry and number of elements used.

K-75
Appendix K

Both models are subjected to the following loadings at room temperature:

• A concentrated vertical load (FZ) at the top of one of the outside columns.

• A concentrated horizontal load in the plane of the wall (FX) at the top of one of the outside

columns. The stiff members described above distribute this shear load evenly to the tops of

all three columns.

• A concentrated transverse load (FY) on the middle column at floor 81.

The above loadings do not include self-weight. Figure K-7 1 shows the various loadings applied to the

ANSYS model.

Simple supports in the plane and out of the plane of the wall (UX,UY) restrain the tops of the columns.

Simple supports in all three directions restrain the bottoms of the columns. The spandrels at the

extremities of the model are free. See Figure K 71.

^lateral (100 k)

vertical (10 k)

UY supports at top, typ.

17.33 in.T

pout-of-plane

(10 k)

UX,UY,UZ supports at base, typ.

Figure K-71. ANSYS model of prefabricated panel showing loading

and boundary conditions.

K-76
Interim Report on Subsystem Structural Analysis of the WTC Towers

Validation Results

Figures K-72 through K-74 show deflected shapes and indicate the displacement at the points of applied

load for the SAP and ANSYS models. Table K-17 summarizes the differences in reactions and

displacements between the SAP and ANSYS models.

UX = 2.91 in.

UX = 2.91 in.

UX = 2.90 in.

UX = 2.91 in.

At Top of Loaded Column

Ave SAP2000: 2.91 in. ANSYS e.g.: 2.70 in. UX difference: 7.3%

UX = 2.70in.

SAP2000 Deflected Shape ANSYS Deflected Shape

Figure K-72. Deflection of prefabricated panels under 100 kip lateral load.

UY = 0.519 in.

UY = G.520 in

UY = 0.519 in.

UY = 0.520 in.

At Point of Applied Load

Ave SAP2000: 0.520 in. ANSYS e.g.: 0.588 in. UY difference: -13.0%

|UY = 0.588 in.

SAP2000 Deflected Shape ANSYS Deflected Shape

Figure K-73. Deflection of prefabricated panels under 100 kip transverse load.

K-77
Appendix K

UZ = -0.00582 in.

UZ = -0,00517 in.

UZ= -0.00668 in.

UZ = -0.00595 in.

UZ = -0.00616 in.

At Top of Loaded Column

Ave SAP2000: -0.00591 in. ANSYSc.g.: -0.00616 in. UZ difference: -4.2%

SAP2000 Deflected Shape ANSYS Deflected Shape

Figure K-74. Deflection of prefabricated panels under 10 kip vertical load.

TableK-17. Validation results.

Loading Condition

SAP2000/ANSYS Displacements"

Difference Range

Lateral FX RX -2 %to+l % UX: 7%

Transverse FY RY• -6% to +7% UY: -13%

Vertical FZ RZ -1 %to +2 % UZ: -4%

a. Displacements considered at tops of columns for FX and FZ, and at points of

load application for FY.

K.8.7 Ongoing Work on the Exterior Wall Subsystem Model

The ongoing work includes the following:

• Stability of a two-story-high exterior column unbraced at the middle floor.

• Stability of a three-story-high exterior column braced at the top and bottom floor levels only.

• Stability of nine-story-high exterior column (floor 92 to 1 00) unbraced at floors 96 and 97

and subjected to fire scenarios.

Response of exterior wall model to fire scenarios

K-78
Interim Report on Subsystem Structural Analysis of the WTC Towers

K.9 FLOOR TRUSS DYNAMIC RESPONSE DUE TO IMPACT OF DROPPING

FLOOR

K.9.1 Impact of Dropping Floor

The failure of dropping floor may occur due to thermal response and/or additional debris weight on the

truss, and/or as a result of the aircraft impact. A floor truss or a group of floor trusses could lose support

at both the exterior and interior supporting ends and drop onto the floor below. This failure mode, which

is shown in Fig. K-75, will be referred to as "full truss drop." Alternatively, a floor truss or a group of

floor trusses could lose support on one side and drop down to impact the floor below. This failure mode,

which is also shown in Fig. K-75, will be referred to as "partial truss drop."

Schematic of full truss drop Schematic of partial truss drop

Before impact

After impact After impact Figure K-75. Schematic of full truss or partial truss drop

and diagonal crushing at impact.

K.9. 2 Purpose and Scope

The purpose of this study is to determine the dynamic response of the target truss from the impact of ftiU and partial truss drop, to determine whether the target truss seats can resist such an impact load and to determine whether the target truss will lose its composite action, become a catenary, and thus fail to restrain the exterior column to which it is connected against instability.

K.9.3 Method of Analysis

The simulation of a floor drop is idealized with a truss drop. This has the inherent assumption that all seats for the floor fail simultaneously to cause a full or partial drop. The dynamic response of the target

truss from the impact of a dropping truss is calculated using conservation of energy principle. The

potential energy of the truss just before drop, which is a function of drop height, converts to the kinetic

energy of the truss just before impact. As the dropping truss starts to impact the target truss, the diagonal

members of the dropping truss are assumed to deform plastically to absorb some of the kinetic energy.

K-79
Appendix K

The energy absorption due to crushing of the furniture and partitions are neglected in this study. The

energy absorption due to diagonal member crushing reduces the kinetic energy available at impact to deform the target truss. All the diagonal members are assumed to deform plastically for the full truss drop case, while only one quarter of the diagonal member are assumed to deform plastically for partial truss drop, representing one quarter of the length of the truss that may come in contact at impact with

floor below. The kinetic energy loss at the time of impact of the dropping truss and the target truss is calculated based on conservation of momentum. The two trusses are assumed to travel together after the

impact, at one-half of the velocity of the dropping floor before impact.

The dynamic load due to the impact of the dropping truss onto the target truss will result in the target

truss to deform plastically beyond the static load due to the weight of the two trusses. The maximum

dynamic deformation of the trusses is calculated by conservation of energy principle assuming that the

resistance of the truss is a bilinear function of displacement. This assumption is based on fitting the FEA

calculated acceleration-deflection relationship of target truss as shown in Fig. K-76.

K.9.4 Results

The ratios of demand-to-seat capacity for the gravity loads of the dropped and impacted trusses moving

together for temperatures of 20 °C, 400 °C, 600 °C, and 700 "C; and the gravity plus dynamic impact loads

for temperatures of 20 °C and 400 °C, are calculated. The demand-to-capacity ratio of less than one shows that the truss seat has sufficient capacity to resist the load, and the demand-to-capacity ratio of

larger than one, implies that the seat could fail. The range of the demand-to-capacity ratios are due to the

different assumptions for the amount of energy loss due to crushing of the diagonal members of the

dropped truss.

The demand-to-capacity ratio of the long-span truss for gravity loads is shown in Table K-18 and for gravity plus impact load is shown in Table K-19. The result for gravity load alone shows that both the

exterior and interior truss seats have sufficient capacity to support the weight of two floors for all temperatures considered. The result for gravity plus impact load shows that at temperatures below 400 °C

neither the exterior nor interior truss seat is expected to fail. Peak deflection response due to gravity and

the dynamic impact of the dropping truss is given in Table K-20. The results show that at room

temperature, and more so at 400 °C, the impacted truss will deflect to an extent that it loses composite

action, and become a catenary. At 400 °C the truss walks off the interior seat. Obviously, a catenary truss

is not able to restrain the exterior column against transverse movement and cannot restrain it from

instability. Although a truss response to increasing acceleration at 700 °C has not yet been developed, the

strength reduction of the truss seats clearly indicates that the failure of truss seats will occur. The results for long-span truss, for partial truss drop, and for the short-span truss are in progress.

K-80
Interim Report on Subsystem Structural Analysis of the WTC Towers

Finite element results Assumed resistance

Knuckle No. 1 to 6 and No. 13 fail, and

truss walks off the interior seat

Bolts at the interior seat shears off

Diagonals start to buckle and 2 knuckles from the interior end fail

20 40 60

Truss mid-span defection (in.)

80

Figure K-76. Target truss resistance against increasing acceleration.

Temp.

(°C)

Demand

(kip)

Capacity (kip) Demand/Capacity

Int. Seat Ext. Seat Int. Seat Ext. Seat 20 26.4 187.3 140.0 0.14 0.19 400 26.4 166.9 125.7 0.16 0.21 600 26.4 81.6 77.8 0.32 0.34 700 26.4 37.2 35.5 0.71 0.74

Table K-19. Demand-to-Capacity ratio of long-span truss

Temp.

(°C) Demand (kip)

Capacity (kip) Demand / Capacity

Int. Seat Ext. Seat Int. Seat Ext. Seat 20 38.6 65.3 187.3 140.0 0.21 0.35 0.28 0.47 400 39.1 45.2 166.9 125.7 0.23 0.27 0.31 0.36

K-81
Appendix K

Table K-20. Peak deflection response due to

static gravity and dynamic Impact.

Temp. (°C)

Static Deflection (in.) Dynamic

Deflection (in.) 20 2.3 7.6 25.4 400 24.2 66.4 89.6

K.9.5 Conclusions

At room temperature, the impact of a dropping truss will not cause failure of truss seats, but will cause the

impacted truss to deform into a catenary. At 400 °C, the impacted truss will walk off the interior seat. In

either case, the impacted floor will not restrain the exterior column against transverse movement and

instability. The impact of a dropping truss at 700 °C will cause failure of truss seats.

K.10 REFERENCES

ACI (American Concrete Institute) Committee 209. 1992. "Prediction of Creep, Shrinkage, and

Temperature Effects in Concrete Structures", ACI 209R-92.

AISC (American Institute for Steel Construction). 2001. Manual of Steel Construction: Load and

Resistance Factor Design, 3rd ed.

AISC (American Institute for Steel Construction). 2003. Manual of Steel Construction, Load and

Resistance Factor Design.

ANSYS, Inc. 2004. Theory Reference, ANSYS, Inc. ASTM (ASTM International). 2001 . "Standard Test Methods for Fire Tests of Building Construction

and Materials," ASTM Standard Test Method El 19.

Bartlett, F.M. and MacGregor, J.G. 1996. "Statistical Analysis of the Compressive Strength of Concrete

in Structures," ACI Materials Journal, March-April, pp. 158-168.

Fields, B.A. and Fields, R.J. January 1991. The Prediction of Elevated Temperature Deformation of

Structural Steel Under Anisothermal Conditions, NIST.

Fields, R. March 22, 2004. Recommended Multiaxial Fracture Criterion, NIST.

LERA (Leslie E. Robertson Associates). 2001. The World Trade Center Design Criteria. Worthington,

Skilling, Helle, and Jackson, original document dated 8/20/65.

LERA (Leslie E. Robertson Associates). August 2003. Development of Structural Databases and

Baseline Models for the World Trade Center (WTC) Towers-Progress Report #4: Structural Analysis

Computer Models.

Phan, Long T. December 1996. Fire Performance of High-Strength Concrete: A Report of the

State-of-the-Art, NIST.

K-82
Interim Report on Subsystem Structural Analysis of the WTC Towers

Phan, Long T. December 2003. Properties of Concrete at Elevated Temperature, NIST.

Seanz, L.P. 1964. "Discussion of Equation of the Stress-Strain Curve of Concrete by Desayi and

Krishman," Journal of the American Concrete Institute, Col. 61, No. 9, pp. 1229-1235.

SHCR (Skilling, Helle, Christiansen, Robertson). 1973a. World Trade Center Drawing Book 8, Concrete

Reinforcing, Chapters 1-4.

SHCR (Skilling, Helle, Christiansen, Robertson). 1973b. World Trade Center Drawing Book 7, Floor

Panels and Truss Details, Chapters 1-5.

Stevick, G.R. March 1994. "Failure of Welds at Elevated Temperatures," Welded Research Council

Bulletin 390, Welding Research Council, pp. 1-39.

K-83
Appendix K

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K-84
Table of Contents

List of Figures L-ii

List of Tables ' L-iv

Appendix L Interim Report on WTC 7 L-1

L.l Building Description L-1

L.1.1 Purpose L-1

L.l. 2 Scope of Work L-1

L. 1 . 3 Introduction L-1

L. 1 .4 Foundations L-3

L. 1 . 5 Con Edison Substation L-4

L.l. 6 Floor Systems L-5

L.l. 7 Columns L-13

L.l. 8 Column Transfer Trusses and Girders L-14

L . 1 . 9 Lateral System L- 1

5

L.2 Observations of Structural Collapse L-1

7

L.2.1 Damage from WTC 1 and WTC 2 Collapses L-17

L.2. 2 Observed Fire Locations L-22

L.2. 3 WTC 7 Collapse Observations L-26

L.2. 4 Interpretation of Collapse Initiation Observations in Elevation L-29

L.2. 5 Interpretation of Collapse Progression Observations L-30

L.2. 6 Debris Field L-33

L.2. 7 Summary L-33

L.3 Collapse Hypothesis L-34

L.3.1 Introduction L-34

L.3. 2 Collapse Initiation Scenarios L-36

L.3. 3 Vertical Progression Scenarios L-41

L.3. 4 Horizontal Progression Scenarios L-46

L.3. 5 Summary of Working Collapse Hypothesis L-50

L.3. 6 Technical Approach for Analysis of the Working Collapse Hypothesis L-52

L.4 References L-52

L-i
List of Figures

Figure L-l. WTC complex L-2

Figure L-2. Size comparison of WTC 7 L-2

Figure L-3. WTC 7 to foundations L-4

Figure L^. Con Ed substation location relative to the WTC 7 building L-5

Figure L-5. Floors 8 to 45 plan L-6

Figure L-6. Floor 1 plan L-8

Figure L-7. Floor 2 plan L-8

Figure L-8. Floor 3 plan L-9

Figure L-9. Floor 4 plan L-9

Figure L-10. Floor 5 plan L-10

Figure L-11. Floor 5 diaphragm plan L-10

Figure L-l 2. Floor 6 plan L-l 1

Figure L-1 3 . Floor 7 plan L-11

Figure L-14. Floor 46 plan L-12

Figure L-15. Floor 47 plan L-12

Figure L-l 6. Roof layout L-l 3

Figure L-l 7. Typical built-up column details L-14

Figure L-l 8. 3D schematic view of transfer trusses and girders between Floors 5 and 7 L-l 5

Figure L-19. Perimeter lateral system elevations L-16

Figure L-20. Core lateral system L-16

Figure L-21. Photograph of roof after WTC 1 collapse L-19

Figure L-22a. Debris damage around Floor 18 of the southwest comer L-20

Figure L-23a. Pedestrian bridge and debris on Vesey Street after WTC 1 collapsed L-22

Figure L-24a. Fires on Floors 1 1 and 12 on the east face L-25

Figure L-25. East penthouse kink L-27

Figure L-26. East penthouse sinks (2.2 s) L-28

Figure L-27. Center screenwall and west penthouse sink (7.9 s) L-28

Figure L-28. Global collapse L-29

Figure L-29. Plan view of regions for collapse initiation L-3

Figure L-30. Likely region of initiating component failures based on videographic and

photographic records of fire and collapse L-31

Figure L-3 1 . Plan View of Collapse Progression L-32

L-ii
Interim Report on WTC 7

Figure L-32. Aerial view of WTC 7 after collapse L-34

Figure L-33. Collapse initiation and vertical progression on the east side of WTC 7 L-35

Figure L-34. Horizontal progression to the west side of WTC 7 L-35

Figure L-35. Phases of WTC 7 building collapse L-36

Figure L-36. Collapse initiation scenarios L-37

Figure L-37. Column 79 capacity versus temperature and unbraced length K L 39

Figure L-38. Effects of temperature gradient on interior column 79 L^O

Figure L-3 9 . Steel strength versus temperature L-4

1

Figure L-40. Vertical collapse progression scenarios L-42

Figure L-^1 . Geometry changes for removal of column 73 L-43

Figure L^2. Geometry changes for removal of columns 70 and 67 L^3

Figure L^3 . Geometry changes for removal of columns 64 and 61 L^4

Figure L^4. Geometry changes for removal of column 76 L^5

Figure L^5. Geometry change for removal of column 79 L-46

Figure L^6. Horizontal collapse progression scenarios L^7

Figure L^7. Transfer components between Floors 5 and 7 L^9

Figure L^8. Horizontal progression mechanism for truss #1 failure L^9

Figure L^9. Horizontal progression mechanism for truss #2 failure L-50

L-iii
List of Tables

Table L-1 . Timeline of WTC 7 collapse as observed from the northwest L-27

L-iv
Appendix L

Interim Report on WTC 7

L.1 BUILDING DESCRIPTION

L.1.1 Purpose

Project 6 addresses the first primary objective of the technical investigation led by the National Institute of Standards and Technology (NIST) of the 47-story World Trade Center (WTC) disaster: to determine

why and how WTC 7 collapsed. Specifically, the objective of this Project is to determine the response of

structural components and systems to the impact damage and fire environment in WTC 7, and to identify

probable structural collapse mechanisms.

L.1.2 Scope of Work

The structural response of WTC 7 to damage from debris and fires is being evaluated to identity possible

collapse sequences and critical components that are consistent with the videographic and photographic

records, interview accounts by individuals that were in or around WTC 7, and other available data. This

work is being conducted in two tasks:

• Task 1 , Structural response analysis to identify critical components

• Task 2, Structural analysis of possible collapse initiation hypotheses

The analytical work is being conducted with the assistance of Gilsanz Murray Steficek LLP.

The scope of work under Task 1 includes (a) develop a nonlinear global structural model of WTC 7 and

evaluate its performance under design gravity loads, (b) identify credible failure sequences for the

structural model with service loads and initial structural damage by analyzing the effect of component

failures (that may have occurred directly or indirectly from fires) on the structural system stability,

(c) identify dominant failure modes for critical components and subsystems determined in (b) for service

loads and elevated structural temperatures, (d) conduct parametric studies of critical subsystems to identify influential parameters, and (e) develop approaches to simplily structural analyses for global

modeling and analyses.

Selected technical results and finding for progress on Task 1 (a), (b), and (c) are presented in the following sections: a description of the WTC 7 strucmral design, observations of damage, fires, and the

structural collapse, and the working collapse hypothesis developed to date.

L.1.3 Introduction

WTC 7 was a 47 story commercial office building, completed in 1987. Its location relative to the WTC

Plaza is shown in Fig. L-1 . It contained approximately 2 million ft' of floor area. The overall dimensions of WTC 7 were approximately 330 ft long, 140 ft wide, and 610 ft tall. The typical floor was

L-1
Appendix L

similar in size to a football or soccer field (see Fig. L-2). The gross floor area was about 75 percent of

that contained in the Empire State Building. The building was constructed over a pre-existing electrical substation owned by Con Edison. The original plans for the Con Ed Substation included supporting a

high-rise building, and the foundation was sized for the planned structure. However, the final design for WTC 7 had a larger footprint than originally planned. Section L.l .4 describes the WTC 7 foundation.

Figure L-1. WTC complex.

A. Comparison to Football Field B. Comparison to Soccer Field

Figure L-2. Size comparison of WTC 7.

L-2
Interim Report on WTC 7

WTC 7 was located immediately to the north of the main WTC Complex, approximately 350 ft from the north side of WTC 1 . It occupied the block bounded by Vesey Street on the south, Barclay Street on the

north, Washington Street on the west, and West Broadway on the east. It was connected to the WTC

complex with a 120 ft wide elevated plaza at the Floor 3, and a 22 ft wide pedestrian bridge, also at Floor 3.

Above Floor 7, the building had typical steel framing for high-rise construction. The floor systems had

composite construction with steel beams supporting concrete slabs on metal deck, with a floor thickness

of 5.5 in. The core and perimeter columns supported the floor system and carried their loads to the foundation. The perimeter moment frame also resisted wind forces. Columns above Floor 7 did not align

with the foundation columns, so braced frames, transfer trusses, and transfer girders were used to transfer loads between these column systems, primarily between Floors 5 and 7. Floors 5 and 7 were heavily

reinforced concrete slabs on metal decks, with thicknesses of 14 in. and 8 in., respectively. The following

sections describe the components and subsystems of WTC 7.

L.1.4 Foundations

WTC 7 and the electrical substation were supported on caisson foundations. When the substation was

constructed in 1967, provision was made for a future office tower by including capacity to carry both the substation and the weight of a future building. Caissons were also installed in the property adjacent to the

substation, for the proposed future building. When WTC 7 was constructed approximately 20 years later,

it was significantly larger than the originally proposed building, and required additional caissons to be

installed, as shown in Fig. L-3.

The typical caisson consisted of several components: a 30-in., 36-in., or 42-in. diameter steel casing, a heavy rolled or built-up steel core shape, vertical reinforcing bars, spiral rebar, and concrete fill. At the base of the caisson core, a pattern of shear studs was placed to help transfer the load from the steel caisson core into the encompassing concrete, from which it passed into the rock. The caissons extended

through the soil, and were socketed (seated) in the bedrock, approximately 60 ft below the surface. There were vertical caissons as well as battered (or sloped) caissons to carry the lateral load. Above the

caissons were heavy grillages composed of built up steel girders. Grillages transferred loads between the

building columns and the caissons.

The distance between the caisson grillages and the first floor varied between 8 ft and 30 ft. This region was braced by reinforced concrete walls with thicknesses varying from 1 ft to 2.5 ft. Many of the WTC 7

steel columns were embedded in these walls, and supporting steel braces were made composite by the addition of shear studs along the height of embedment.

Areas between the concrete walls were backfilled with compacted gravel fill and then covered with a concrete slab on grade or framed slab to form closed cells and bring the structure up to the required

elevation. In some cases, the area was left unfilled and used to house fiiel tanks.

L-3
Appendix L

o Indicates additional WTC 7 caisson

^

Indicates original caisson

Source: McAllister 2002.

Figure L-3. WTC 7 to foundations.

L.1.5 Con Edison Substation

The Con Ed Substation was constructed in 1967 and consisted of a steel framed structure with cast-in- place concrete floors and walls. It was placed on the northerly portion of the site and extended

approximately 40 ft north of the north facade of WTC 7, as shown in Fig. L^. Its southerly boundary was irregular, but extended approximately one-third to two-thirds of the width of WTC 7. The Con Ed

Substation was three stories in height.

The substation's lateral system consisted of a moment frame along the northern row of interior columns.

Along the south edge of the substation there was a braced frame. This braced frame was coincident with

the north side of the WTC 7 core, at columns 64, 67, 70, and 73. Lateral loads from WTC 7 were passed

directly from the core above to the Con Ed braced frame below. There were also two moment frames

within the substation oriented in the north-south direction, one on each end of the WTC 7 core.

The WTC 7 columns, which were within the perimeter of the substation, were supported by substation

columns. During the construction of WTC 7, heavy plates were welded to the tops of the existing

substation columns which then received the new building columns.
Interim Report on WTC 7

L.1.6 Floor Systems

Typical Floor Systems Above Floor 7

The typical floor framing system, shown in Fig. L-5, was composed of rolled steel wide-flange beams

with composite metal decking and concrete slabs. Floors 8 through 45 had essentially the same framing

plan, but the core layout varied over the height of the building.

Floors 8 through 45 had floor slabs that were composed of 3 in., 20 gage metal deck with 2.5 in., 3,500 psi normal weight concrete, for a total thickness of 5.5 in. There was one layer of 6x6 W1.4xW1.4

welded wire mesh within the concrete. The drawings show a second layer of mesh placed over girders at the slab edges. The fastening requirements for the metal deck are not shown on the drawings, but standard

practice provides puddle welds 12 in. on-center at the beams and side lap welds, screws, or buttonpunching at 36 in. on-center between adjacent panels of deck. The drawings contain a note calling for

1.5 in., 20 gage deck with 4 in. concrete topping (5.5 in. total) in the elevator lobbies, where there was a

3 in. floor finish specified by the architect.

Typical floor framing for Floors 8 through 20 and Floors 24 through 45 consisted of 50 ksi wide-flange

beams and girders. Between the core columns was a grid of beams and girders. Core girders ranged in size fi-om W16x31 to W36xl35, depending on the span and load. (W 16x31 describes a steel wide-flange

beam, sometimes referred to as T' beams; the nomenclature indicates the cross-section is nominally 16 in. deep and weighs 31 lb per lineal foot.) Beams spanned directly between the core and the exterior of the

building, at approximately 9 ft on-center spacing. On the north and east sides, the typical beam was a W24x55 with 28 shear studs, spanning 53 ft. On the south side, the typical beam was a W16x26 with

24 shear studs sparming 36ft. Between the exterior columns were moment connected girders that formed

part of the lateral system of the building. On Floors 10, 19, and 20, a portion of the floor framing was

L-5
Appendix L

Number Beams

Figure L-5. Floors 8 to 45 plan.

reinforced with plates attached to the bottom flange. Certain connections at these floors were also reinforced.

Floors 21 to 23 had slightly heavier steel framing than the typical floors. Core girders were generally one

size class larger than the typical floor; the beams between the core and the south facade were W16x31

instead of W16x26. There were additional studs on the W24x55 beams on the north and west sides.

Most of the beams and girders were made composite with the slabs through the use of shear studs. Typically, the shear studs were 0.75 in. in diameter by 5 in. long, spaced 1 ft to 2 ft on center. Studs were

L-6
Interim Report on WTC 7

not indicated on the design drawings for many of the core girders. The design drawings specified design

forces for connections and suggested a typical detail, but did not show specific connection designs; this is standard practice on the U.S. east coast. The erection drawings indicate that design shear forces for the

typical beam and girder connections were to be taken from the American Institute of Steel Construction

(AISC) beam design tables for beams without shear studs, using 1.5 times those forces for beams with

shear studs.

According to a paper by Salvarinas (1986), who was the project manager for Frankel Steel, which

fabricated the steel for WTC 7, the typical floor beam to girder and girder to core column connection was

a single shear plate, although end plate and double angle connections were also used. The typical beam to exterior column connection was a seated connection. The typical bolt size for the simple shear

connections is cited as 0.875 in. in diameter ASTM A325, where A325 is a standard specification for a

structural bolt specified by ASTM International. The bolt size used for heavier brace and moment

connections was 1 in. in diameter ASTM A490. Information on the specific connection details used is unavailable at this time.

Other Floors

The remaining floors. Floors 1 to 7 and Floors 46 to 47, were atypical and are described below and in Figs. L-6 through L-15.

Floor 1 was built adjacent to the substation and included the truck ramp for the WTC complex. The first floor is shown in Fig. L-6. The floor was framed with steel beams that were encased in a formed

concrete slab. The floor slab was 14 in. thick, with typical #5 reinforcement bars (5/8 in. rebar) at a 10 in. to 12 in. spacing and #6 rebar at 9 in. spacing for the bottom reinforcement; #5 rebar at 12 in. spacing was used for temperature reinforcement. The southeast portion of the floor above the WTC truck ramp

had a 6 in. formed concrete slab with #4 rebar at 12 in. spacing for top and bottom reinforcement;

#4 rebar at 18 in. spacing was used for temperature reinforcement.

The floor slab for Floors 2, 3, 4, and 6 had a 3 in., 20 gage metal deck with 3 in. 3,500 psi normal weight

concrete, for a total thickness of 6 in. Floors 2 and 3 were also partial floors adjacent to the substation. In addition, they had a floor opening on the south side to form the atrium above the ground level lobby (see

Figs. L-7 and L-8). Floor 4 was above the substation and had a large opening over most of the south side

of the building, to form a double-height space above the 3rd floor lobby (see Fig. L-9). Floor 6 had two

openings on the floor to form a double-height mechanical space, one at the east side and the other one at the southeast comer (see Fig. L-12). Truss #2 and column 80 were located in this double-height

mechanical space.

The 5th floor slab was 1 1 in. of 3,500 psi normal weight concrete on top of a 3 in., 18 gage composite

metal deck for a total slab thickness of 14 in. The slab was heavily reinforced, with #7 rebar at 12 in. spacing for top reinforcement in both directions and #9 rebar at 12 in. spacing for bottom reinforcement

that acted as additional diaphragm chord reinforcement in many areas. This floor also had 36 ksi steel WT sections (W, or wide-flange, sections cut in half to look like a 'T' section) embedded in the 1 1 in. concrete slab above the deck. The WT sections were designed to act as a horizontal truss within the plane

of the floor between the perimeter and core columns (see Figs. L-10 and L-1 1).

L-7
Appendix L

Figure L-6. Floor 1 plan.

Figure L-7. Floor 2 plan.

L-8
Interim Report on WTC 7

Figure L-8. Floor 3 plan.

Figure L-9. Floor 4 plan.

L-9
Appendix L

L-10
Interim Report on WTC 7

Figure L-1 3. Floor 7 plan.

The 7th Floor slab consisted of 5 in. of 3,500 psi normal weight concrete on top of 3 in., 18 gage

composite metal deck, for a total thickness of 8 in. The slab was reinforced with #5 rebars at 6 in. oncenter in both directions. Regions of the slab on the south side of the building had 8 in. of formed

concrete without any metal deck. In these regions two layers of steel reinforcing were provided (see

Fig. L-13).

L-11
Appendix L

Figure L-14. Floor 46 plan.

Figure L-1 5. Floor 47 plan.

Floors 41 and 43 had the east half removed to provide double height spaces. Columns in these areas and

areas of Floors 40 and 42 had been reinforced to provide adequate capacity for the additional height and

change in use by tenants. By 2001, Floors 41 and 43 had been reconstructed to provide full floor space. Specifics of this reconstruction are not available at this time.

The 46th Floor had heavier framing to support the cooling towers and dunnage on the north side,

(alternating W36xl50 with W36x260 under the posts) and the setback roof on the south side (alternating

W21x44 with W36xl50 under the posts). There was a 6 in. reinforced concrete slab in a portion of the core and under the cooling towers (see Fig. L-14).

L-1
Interim Report on WTC 7

Floor 47 had a double height space extending from the 46th Floor to the underside of the roof for the

cooling towers on the north side. There was also a setback roof on the south side at Floor 46

(see Fig. L-15).

Roof and Penthouses

The roof had a concrete slab on metal deck, the top of which sloped 3 in., from an 8.5 in. thickness to a

5.5 in. thickness, to provide drainage. The wire mesh in this slab was 6x6 W2.4xW2.4, which was

70 percent heavier mesh than at the typical floor. There were slab openings for the cooling towers on the north and the setback roof on the south. The area above the cooling towers was framed in steel, with

areas of grating spanning between the beams. A series of diagonal WT 6x9 members under the grating

provided diaphragm action in this area. The east side of the floor was reinforced to carry the east penthouse and its contents. Specifics of this reinforcement are not available at this time.

The west penthouse roof was framed in steel with the floor slab increased to a 6 in. thickness. The

framing and roof reinforcement for the east penthouse and the mechanical equipment screenwall are not

available at this time. Layout of these areas has been determined from photographs, as shown in Fig. L-16.

D n D Q Q 5 5 B 5 5

Cooling Towers

Below

9...S)l.S)....(^....(jk...Q..l&

I ° East

; West

Penthouse Screenwall

,® „0 .0

N

/-\ reninoujst

^

„® M „® 3

Penthouse

,0

,0 3

Roof Setback

J g D g B a b d_

Figure L-16. Roof layout.

L.1.7 Columns

Core columns were primarily rolled wide-flange shapes of grade 36 or 50 steel. As the loads increased

towards the base of the building, many of these column sizes were increased through the use of built-up

shapes. These built-up columns had a W14x730 core with cover plates welded to the flanges (to form a box) or web plates welded between the flanges as shown in Fig. L-17. The reinforcing plate welds were

L-13
Appendix L

specified to be continuous 0.5 in. fillet welds at the cover plates and 0.3 1 3 in. minimum at the web plates.

Plate thickness ranged from 1.5 in to 8 in. Reinforcing plates were specified as follows:

Plate thickness t (in.):

2 < t < 4 ASTM A588 Grade 50

4 < t < 6 ASTM A5 72 Grade 42

t>6 ASTM A5 88 Grade 42

Figure L-17. Typical built-up column details.

Typical core column splices were shown on available erection drawings. The adjoining surfaces of

columns were specified to be milled. The splice plates were welded or bolted to the outsides of the column web and flanges. Built-up columns were also milled at their bearing ends but the splice plates were fillet welded to the cover plates.

Perimeter columns were nominally 14 in. wide-flange shapes (W14) of ASTM A 36 steel. Perimeter

column splices were similar to the core column splices.

L.1 .8 Column Transfer Trusses and Girders

The layout of the substructure and Con Edison columns did not align with the column layout in the upper

portion of WTC 7. Therefore a series of column transfers were constructed. These transfers occurred

primarily between Floors 5 and 7. See Fig. L-18 for a schematic rendering of the transfers.

Columns 47 through 54, at the north facade, were transferred at Floor 7 by cantilever girders to bring

them in line with the substation columns, offset 6 ft to 9 ft to the south. The back-span of these

cantilevers was supported by the north side core columns. The eastern most cantilever girder was

connected to truss #1, and the western most cantilever girder was connected to truss #3 (see Fig. L-18).

L-14
Interim Report on WTC 7

Figure L-18. 3D schematic view of transfer trusses and girders between Floors 5 and 7.

Column 76 was supported at Floor 7 by truss #1 . The west side of truss #1 is supported by column 73,

while the east side is supported by a transfer girder running north-south which is, in turn, supported by

columns E3 and E4 at Floor 5.

Columns 58, 59, and 78 were transferred by simply supported girders at Floor 7. Column 78 was

supported at Floor 7 by a transfer girder that was supported at its north end by truss #2. Column 77 was

also supported by truss #2. Truss #2 was supported by column 74 at its west end and by column 80 at its east end.

Column 61 was supported by truss #3. Truss #3 runs north-south and was supported by columns 62 and

61 A. Truss #3 has a 10 ft cantilever span between column 61 and column 61 A and an 18 ft back span to column 62.

L.1.9 Lateral System

Above Floor 7, WTC 7 had a perimeter moment frame. Exterior columns were typically rolled W14

shapes of ASTM A36 grade steel. Column trees were fabricated for the east and west facades with field splices occurring every other story in the columns and at the spandrel beam midspan between columns,

where the tree stubs were spliced with a bolted connection. On the north and south facades, the moment

frames were constructed with spandrel connections at the face of the columns. Some column splices were shown on the erection drawings to be partial penetration groove welds between the column flanges.

L-15
Appendix L

At Floors 5 to 7 and Floors 22 to 24, there was a perimeter belt truss, shown in Fig. L-19. Below Floor 7, a combination of moment and braced frames around the perimeter and a series of braced frames in the

core, is shown in Fig. L-20. The strong diaphragms of Floors 5 and 7 transferred load from the perimeter

to the core. Above the loading dock at the south facade, two of the columns were hung from the belt truss

at Floors 5 through 7. Above the Con Edison vault at the north facade, eight columns were also hanging

from the belt truss between Floors 5 and 7.

w w

Con Ed

Substation

Shipping

Ramp

Shipping - Ramp

NORTH SOUTH WEST EAST

Figure L-19. Perimeter lateral system elevations.

NORTH CORE SOUTH CORE TRANSVERSE CORE

Figure L-20. Core lateral system.

L-16
Interim Report on WTC 7

L.2 OBSERVATIONS OF STRUCTURAL COLLAPSE

This section presents observed data and events from available drawings, photographic and videographic

records, interviews, and other data sources for WTC 7 to identify damage and fire locations. Damage to WTC 7 from debris impact from WTC 1 and WTC 2 is summarized in Section L.2. 2, followed by known

fire growth and progression in Section L.2. 3. The observed exterior sequence of collapse events from

photographic and videographic records are described in Sections L.2. 4 and L.2. 5, where collapse

observations are considered from the plan and elevation views of the structure, respectively. These

observations have been used for developing possible collapse initiation locations and progression

mechanisms, which are presented in Section L.3.

L.2.1 Damage from WTC 1 and WTC 2 Collapses

To place the events leading to the global collapse of WTC 7 into context, it is helpful to summarize the events of September 1 1 , 200 1

8:46 a.m. WTC 1 was struck by an aircraft 9:03 a.m. WTC 2 was struck by an aircraft 9:59 a.m. WTC 2 collapsed

10:28 a.m. WTC 1 collapsed

5:21 p.m. WTC 7 collapsed

After WTC 1 collapsed, the south face of WTC 7 was obscured by smoke, making direct observation of

damage from photographs or videos difficult or impossible. The source of the smoke is uncertain, as

large fires were burning in WTC 5 and WTC 6, as well as those noted below in WTC 7. The light but

prevalent winds from the northwest caused the smoke to rise on the leeward, or south, side of the

building. The following information about damage seen in WTC 7 was obtained from interviews of

people in or near the building:

After WTC 2 collapsed:

• Some south face glass panes were broken at lower lobby floors

• Dust covered the lobby areas at Floors 1 and 3

• Power was on in the building and phones were working

• No fires were observed

Reported close to time of WTC 1 collapse:

• East stair experienced an air pressure burst, filled with dust/smoke, lost lights

• West stair filled with dust/smoke, lost lights, swayed at Floors 29 through 30, and a crack was

felt (in the dark) on the stairwell wall between Floors 27 through 28 and Floors 29 through 30

• Floors 7 and 8 had no power, air was breathable but not clear

L-17
Appendix L

• Phone lights on Floor 7 were on but could not call out

After WTC 1 collapsed:

• Heavy debris (exterior panels fi-om WTC 1) was seen on Vesey Street and the WTC 7 promenade structure at the third floor level

• Southwest comer damage extended over Floors 8 to 1

8

• Damage was observed on the south face that starts at the roof level and severed the spandrels

between exterior columns near the southwest comer for at least 5 to 10 floors. However, the extent and details of this damage have not yet been discemed, as smoke is present.

• Damage to the south face was described by a number of individuals. While the accounts are mostly consistent, there are some conflicting descriptions:

- middle one-fourth to one-third width of the south face was gouged out from Floor 10 to the ground

- large debris hole near center of the south face around Floor 14

- debris damage across one-fourth width of the south face, starting several floors above the atrium (extended from the ground to 5th floor), noted that the atrium glass was still intact

- from inside the building at the 8th or 9th Floor elevator lobby, where two elevator cars were ejected from their shafts and landed in the hallway north of the elevator shaft, the

visible portion of the south wall was gone with more light visible from the west side

possibly indicating damage extending to the west

At 12:10 to 12:15 p.m.:

• Firefighters found individuals on Floors 7 and 8 and led them out of the building

• No fires, heavy dust or smoke were reported as they left Floor 8

• Cubicle fire was seen along west wall on Floor 7 just before leaving

• No heavy debris was observed in the lobby area as the building was exited, primarily white

dust coating and black wires hanging from ceiling areas were observed

Photographs support some of these reports and show additional damage at the upper portions of the building. Figure L-21 is an aerial view of WTC 7 after the collapse of WTC 1. There is no visible debris on the roof; some minor damage is seen on the south side at the parapet wall. Figures L-22a and L-22b

show the reported damage between Floors 8 to 1 8 at the southwest comer. Much of the damage above

Floor 18 appears to be nonstmctural. The black areas on the facade indicate areas of burned out fires. Note the heavy smoke obstracting any observations along the south face. Study of this photograph

indicates that at least two exterior columns were severed. Figures L-23a and L-23b show the debris on Vesey Street in front of WTC 7 after the collapse of WTC 1 . The pedestrian bridge (L-23a) and the

L-18
Interim Report on WTC 7

promenade (L-23b) appear to be standing, although damaged. Exterior panels from WTC 1 can be seen on Vesey Street and on the promenade. The approximate extent of possible damage due to debris from

WTC 1 is shown in Fig. L-23c.

Figure L-21. Photograph of roof after WTC 1 collapse.

L-19
Appendix L

A.. ©2001. New York City Police Department. All rights reserved.

^

SW Corner

Damage

Starting at Floor 18

Figure L-22a. Debris damage around Floor 18 of the southwest corner.

L-20
Interim Report on WTC 7

Figure L-22b. Debris damage around Floor 8 of the southwest corner.

L-21
Appendix L

Figure L-23a. Pedestrian bridge and debris on Vesey Street after WTC 1 collapsed.

L.2.2 Observed Fire Locations

Photographs and videos were used to determine fire locations and movement within WTC 7. Most of the

available information is for the north and east faces of WTC 7. Information about fires in other areas of

the building was obtained from interviews, and is summarized as follows:

From 1 1 :30 a.m. to 2:30 p.m.:

• No diesel smells reported from the exterior, stairwells, or lobby areas

• No signs of fire or smoke were reported below the 6th Floor from the exterior, stairwells or lobby areas

• In the east stairwell, smoke was observed around Floors 1 9 or 20, and a signs of a fully

involved fire on the south side of Floor 23 were heard/seen/smelled from Floor 22.

• Interviews place a fire on Floor 7 at the west wall, toward the south side, at approximately

12:15 p.m.

L-22
Interim Report on WTC 7

Figure L-23b. WTC 7 Promenade and debris on Vesey Street after WTC 1 collapsed.

Southwest Corner Possible Roof and Possible Region of Impact

Debris Damage Upper Level Damage Damage by WTC1 Debris Figure L-23c. Possible extent of debris damage in plan.

L-23
Appendix L

• From West and Vesey Streets near the Verizon Building, fires were observed in floors estimated to be numbered in the 20s and 30s.

Looking from the southwest comer at the south face:

• Fire was seen in the southwest comer near Floor 10 or 11

• Fire was seen on Floors 6, 7, 8, 2 1 , and 30

• Heavy black smoke came out of a large, multi-story gash in the south face

Looking from the southeast comer of the south face:

• Fire seen on Floor 14 (reported floor number) on south face; the face above the fire was

covered with smoke

• Fire on Floor 14 moved towards the east face

Looking at the east face:

• Fire on Floor 14 (reported floor) moved along east face toward the north side

Photographs and videos were used with these interview accounts to document fire progression in the

building. The fires seen in photographs and videos are summarized:

Before 2:00 p.m.

• Figures L-22a shows fires that had bumed out by early aftemoon on Floors 19, 21, 22, 29,

and 30 along the west face near the southwest comer.

2:00 to 2:30 p.m.

• Figure L-24a shows fires on east face Floors 1 1 and 12 at the southeast comer. Several

photos during this time show fires progressing north.

3:00 to 5:00 p.m.

• Around 3 p.m., fires were observed on Floors 7 and 12 along the north face. The fire on

Floor 1 2 appeared to bypass the northeast comer and was first observed at a point

approximately one third of the width from the northeast comer, and then spread both east and

west across the north face.

• Some time later, fires were observed on Floors 8 and 13, with the fire on Floor 8 moving

from west to east and the fire on Floor 1 3 moving from east to west. Figure L-24b shows

fires on Floors 7 and 12.

• At this time, the fire on Floor 7 appeared to have stopped progressing near the middle of the

north face.

L-24
Interim Report on WTC 7

L-25
Appendix L

Figure L-24b. Fires on Floors 7 and 12 on the north face.

• The fire on Floor 8 continued to move east on the north face, eventually reaching the

northeast comer and moving to the east face.

• Around 4:45 p.m., a photograph showed fires Floors 7, 8, 9, and 1 1 near the middle of the north face; Floor 12 was burned out by this time.

L.2.3 WTC 7 Collapse Observations

The collapse of WTC 7 was recorded on several videos from locations northeast and northwest of the building. Study of these videos led to the development of the timeline in Table L-1, which lists the

visible external sequence of events. Figures L-25 to L-28 are images from a CBS News Archives video

that show key points observed during the collapse.

The deformed shape of the east penthouse roof shows that the middle fell before the sides (see Fig. L-25),

as the whole penthouse drops into the main building (see Fig. L-26). This may imply that support

initially remained on the east and west edges of the east penthouse. Therefore, the perimeter columns on

L-26
Interim Report on WTC 7

the east side of the building which have not aheady been considered least likely, may be considered less likely locations for collapse initiation.

Table L-1. Timeline of WTC 7 collapse as observed from the northwest.

Time

Interval

(s)

Total

Time (s) Observation from CBS Video

u.u u.u - rirsi muvcmciii oi cdsi pciiinousc looiiinc Qownwarus

0.9 0.9 - East penthouse kink between columns 44 and 45 (Fig. L-25)

- First 2 windows at Floor 40 fail between columns 44—45 (windows 9 and

1 1 from east end)

0.3 1.2 - 4 more windows fail at Floor 40

- East penthouse submerged from view (now inside building)

0.5 1.7 - 3 windows break at Floor 41, Floor 43, Floor 44

0.5 2.2 - East penthouse completely submerged (Fig. L-26)

1 0

1 .0 4.U - Windows break along column 46 at Floors 37 and 40

3.0 7.0 - West penthouse and screenwall begin to move downward into building

- Movement of entire north face of WTC7 (visible above Floor 2 1

)

0.2 7.2 - West end of roof starts to move

0.5 7.7 - East end of roof starts to move

- Kink formed in north facade along column 46-47

0.4 8.2 - West penthouse and screenwall submerged

- Windows fail between Floors 33-39 around column 55

- Global collapse initiates (Fig. L-27 and L-28)

East Screen- West

Penthouse wall Penthouse

Figure L-25. East penthouse kink.

L-27
Appendix L

East Penthouse

Sinks into

Building

Figure L-26. East penthouse sinks (2.2 s).

West Penthouse

© 2001 CBS News Archives,

Figure L-27. Center screenwall and west penthouse sink (7.9 s).

L-28
Interim Report on WTC 7

North

Facade Kink

Figure L-28. Global collapse.

Possible Locations of Collapse Initiation

Columns 76, 77, 78, 79, 80, and 81 appear to have direct influence on the collapse initiation of the east penthouse. A failure of any of these columns, truss #1 or #2, or the east transfer girder, or some

combination of these components, with possible contribution of adjacent framing and floor systems, could

be considered possible locations of the initiating events that led to the observed collapse of the east penthouse.

L.2.4 Interpretation of Collapse Initiation Observations in Elevation

In addition to determining some possible locations of the collapse initiation locations within the plan of

the structure, it is also helpful to use the available collapse documentation to identify possible locations in the building elevation for the initial failure.

L-29
Appendix L

Least Likely Locations of Collapse Initiation-Penthouse Failure Mechanism

Because the first visible failure is in the east penthouse, one possible collapse initiation mechanism

involves a local failure of the penthouse framing, which then progressed down the structure with floors sequentially impacting upon those below. There are two reasons that this scenario may be considered

unlikely.

First, there was no visible abnormal loading locally applied to cause a local failure at the East Penthouse.

The photograph in Fig. L 2 1 shows the east penthouse sustained no damage due to the collapse of the

WTC towers. The videographic records do not show any visible fire in or near the penthouse prior to collapse.

Second, Fig. L-25 shows a snapshot as the east penthouse starts to collapse. When the roof of the penthouse starts to fall, a line of windows (roughly in line with columns 79 to 81) has broken over the

entire height of the visible region. In free fall, it would take 3 to 4 seconds for an object to fall from the

roof elevation to the height of the bottom visible broken window, around Floor 33. Since the bottom

window is broken nearly simultaneously when the kink is seen at the east penthouse, the initial failure may be assumed to have propagated upward from the lowest window breakage rather than propagated

downward from the top of the building. Therefore, initial failure within the penthouse may be considered

unlikely.

Less Likely Locations of Collapse Initiation-High Elevation Column Failure Mechanism

Another possible collapse initiation mechanism may be the failure of a column in the upper elevations of

the building. The collapse could have progressed vertically upward by pulling down the floors above the

failed column as debris landed on and sequentially crushed the floors below.

The timing required for this mechanism, in accordance with gravitational acceleration, requires that any

column locations significantly above the 13th floor (the lowest visible floor in photographic and

videographic records) may be considered unlikely failure initiation locations. The lack of observed fires in the floors above Floor 1 3 also reduces the likelihood of failure initiating in this region of the building.

Possible Locations of Collapse Initiation Mechanism

Based on review of the photographic and videographic records, a failure of any column within the plan

area shown in Fig. L-29, and below Floor 13, likely contributed to the collapse initiation. This includes

columns 76, 77, 78, 79, 80, and 81, truss #1, truss #2, column 78A, the east transfer girder and adjacent

framing and floor systems within this region (see Fig. L-30).

L.2.5 Interpretation of Collapse Progression Observations

Interior columns 79, 80, and 8 1 , were located directly below the east penthouse on the roof and supported

large tributary areas. It appears that some sequence of component failures in the region identified in Figs. L-29 and L-30 led to the failure of one or more of these columns, as discussed above. The failure progressed vertically upward within the failed bay to the roof level, based upon observations of window

breakage relative to failure of rooftop structures, and was first visible from the exterior when the east penthouse lost support (see Fig. L-26).

L-30
Interim Report on WTC 7

Columns Designation

^1 Possible Location of Initial Failure

I I Less Likely Location of Initial Failure

I I Least Likely

I I Approximate Location Approximate Location of of First kink in East Window Breakage Below Penthouse

Southwest Corner

Debris Damage

Roof and Upper Approximate Region of Impact

Level Debris Damage by Large WTC1 Debris

Damage

Figure L-29. Plan view of regions for collapse Initiation.

Figure L-30. Likely region of initiating component failures based on videographic

and photographic records of fire and collapse.

L-31
Appendix L

The 5 s to 6 s delay between the failure of the east penthouse and the failure of the screenwall and west

penthouse (shown in Fig. L-27) approximates the time it would take for the debris pile from the vertical

failure progression on the east side of the building to reach Floors 5 to 7 and damage the transfer trusses and girders in this area. A kink developed in the north facade approximately where column 76 projects to the north face. The

kink may have formed in the plane of the north facade or it may represent a displacement in the structure along this line towards the south. The area of this kink correlates to the easternmost cantilever transfer at Floor 7. All of the Floor 7 cantilever transfer girders had back spans supported along the line of the north core columns, of which the easternmost one was supported by truss # 1. This north facade kink also coincides with the girders at the eastern edge of the cooling tower area at Floor 46.

When the screenwall and the west penthouse sank into the building, a line of windows broke from

Floor 44 down to the bottom of the visible range, which is approximately at Floor 33 on the west side of

the structure (see Fig. L-27). This area aligns with column 61, which is supported by the cantilevered

end of transfer truss #3 between Floors 5 and 7, as shown in Fig. L-31. This suggests that the observed

window breakage may be related to the failure of column 61 or truss #3.

Approximate Location

of Vertical line of window breakage

Approximate Location of Kink in East Penthouse

^1 Reoion of Initial Collaose

I I Region of Secondary Collapse

Unknown Time of Collaose i

I I Final Collapse j

Approximate Location of I

North Facade Kink

TRUSS its- Cooling

Towers

»' West

I Penthouse

TRUSS #1

Screenwall TRUSS R #2 'I

Southwest Corner Roof and Upper Level Approximate Region of Impact

Debris Damage Debris Damage Damage by Large WTC1 Debris

Figure L-31. Plan View of Collapse Progression.

L-32
Interim Report on WTC 7

The simultaneous failure of screenwall and west penthouse structures, window breakage on the west side

of the north facade, and initiation of global collapse (see Fig. L-28) indicates that the building loads could no longer be supported. Horizontal progression of the collapse appears to have occurred after the vertical collapse on the east side of the building. The greater strength of Floors 5 and 7 relative to the other floors and the transfer trusses between these floors suggests that this region of the building played a key role in destabilizing the remaining core columns, and the global collapse occurred with few external signs prior

to the system failure.

All of the photographic and videographic records show the north facade collapsing from below the visible

area; the facade appears to sink into the ground without any sign of the other floors in the visible portion

of the building collapsing. This may indicate that the collapse of the facade starts below the area visible

in the photographic and videographic records.

L.2.6 Debris Field

The debris of WTC 7 was mostly contained within the original footprint of the building. From aerial photos, the debris visible on top of the pile is mostly fa9ade structure. This failure sequence suggests that the interior of the building collapsed before the exterior. See Fig. L-32.

L.2.7 Summary

The possible region of collapse initiation and progression has been refined and can be limited based upon

available data as follows:

• Based upon the observed fire locations, it appears that the initiating collapse event may have

occurred on Floors 5 through 13.

• Due to the pattern of window breakage, it appears that the initiating collapse event may have

occurred below Floor 13 and then progressed vertically upward to the east penthouse.

• Since the middle of the east penthouse roofline appears to fall first, it is possible that the

initiating collapse event occurred at columns or transfer components with direct influence on

the footprint of the east penthouse.

• The north facade kink and the window breakage on the west side of the north facade as the screenwall and west penthouse began to fall into the building core suggest that a horizontal

collapse mechanism occurred between Floors 5 and 7, as there are vertical discontinuities in

line with each of these elements between Floors 5 and 7.

• The relatively small debris field, with the exterior moment frame visible on top of the

building debris, an internal collapse mechanism is likely.

L-33
Appendix L

Figure L-32. Aerial view of WTC 7 after collapse.

The working collapse hypothesis can be summarized in Figs. L-33 and L-34, which illustrate the

components of the observed collapse event: collapse initiation and vertical progression, horizontal

progression, and global collapse.

L.3 COLLAPSE HYPOTHESIS

L.3.1 Introduction

WTC 7 suffered a global collapse. The initiating cause or causes of this collapse, and its sequence of

events, are still being investigated though fire appears to have played a key role and there may have been some physical damage on the south side of the building.

To develop a working hypothesis for the collapse sequence, it is useful to subdivide the problem into several phases. Many factors and structural components may have contributed to the start of the collapse,

but there must have been an initiating event. After the collapse initiated, it progressed to other parts of

the building, leading to their failure as well. From the observations of the collapse (see Section L.2), it appears that fu-st there was a vertical failure progression, from some point in the lower eastern portion of

the building up to the east penthouse. After a time lag of approximately five seconds, the screenwall and

west penthouse were observed to begin sinking into the core area. This suggests that there was a

horizontal progression of the collapse towards the west. Since the screenwall and west penthouse fell almost simultaneously, it is reasonable to assume that the horizontal progression captured all the columns

that support these building parts.

L-34
Interim Report on WTC 7

East Penthouse Screenwall

48th Fl

40'" Fl

30th Fl

Window

Breakage

around

Fl 44

© 20C 1 tef News

13th Fl

7th Fl

5th r

'

Archives,

Vertical Collapse

Progression

I

Failed Column

.1 1 1 i j-ii

Figure L-33. Collapse Initiation and vertical progression on the east side of WTC 7.

East Penthouse

Sinks into Building

48th Fl

40th Fl

30th Fl © 201 II CBS hjews Archives

13th Fl

JOtJi Fl_

_7tji Fl_

~5t'h fT

Substation 1

Debris Impacts

Strong Floor / Truss Below

Horizontal Collapse

Progression

Remainder of Building Sinks

48th Fl

40th Fl

30th Fl ©20_31 CBS ^\ev^^ Arphivbs. Global

J Collapse

III'

Window

Breakage

Figure L-34. Horizontal progression to the west side of WTC 7.

L-35
Appendix L

Within each phase of the collapse shown in Fig. L-35, the initiating event, vertical progression, horizontal

progression, and global collapse, exist many possible scenarios. Scenarios have been developed from

available observations of the collapse and are explored with event trees. Preliminary analyses, combined

with the observations of collapse, can be used to prune the list of postulated scenarios to a relatively small

number of possible collapse hypotheses. These possible hypotheses will then be analyzed in successive

levels of detail to try to determine one or more probable sequence of events leading to the building

collapse.

Initiating Event Vertical Collapse

Progression

Horizontal Collapse

Progression

Global Collapse

Figure L-35. Phases of WTC 7 building collapse.

L.3.2 Collapse Initiation Scenarios

For the collapse to have started, there must have been a component or group of components that failed

first, referred to here as the initiating event, as shown in Fig. L-36. The initiating event may have

included structural components severed or damaged by falling debris (II. 1) and/or structural components

affected by fires (II .2).

II. 1 Initiating Components Fail Due to Debris Damage From WTC 1 of WTC 2: The initiating

components may have included perimeter or interior columns that were severed or damaged by falling

debris from WTC 1 or WTC 2.

• 12.1 Debris Damage to South Facade Columns: Perimeter columns on the south face and

the southwest comer were reported or observed in photographic and videographic records to have been severed or damaged after WTC 1 collapsed. If the initiating event was due to damage to the perimeter moment frame, then it would have started along the south or southwest facade. Photographic and videographic records show that columns on the north

and east facades were undamaged by debris impact.

- 13.1 Perimeter Moment Frame Arrests Failure Progression: Analysis of the global

structure indicates that the structure redistributed loads around the severed and damaged

areas. A progression of column failure to adjacent columns would have been arrested by

the vierendeel action of the perimeter moment frame, which could span across a sizeable

opening due to the strength and stiffness of the frame.

- 12.2 Debris Damage to Interior Columns: Interior columns may have been severed or damaged by impacting debris.

- 13.2 Interior Columns Fail Immediately: If interior columns had been severed or severely deformed, they may have failed immediately.

L-36
Interim Report on WTC 7

11.1 Initiating event components

damaged by- impacting debris

11.2 Initiating event components

fail due to fire effects

13.1 Perimeter

moment frame

redistributes loads around damage

12.1 South facade

damaged

12.2 Interior columns and

adjacent framing

damaged

12.4 Components

on floors 5-8 and

11-13 with fires in afternoon

13.2 Interior columns fail immediately

13.3 Weakened

components more

susceptible to fire effects. Go to 11.2.

13.6 Fire fails interior columns,

east transfer

girder, and/or

transfer truss components

14.1 Local failure only

12.3 Components

on floors with fires that burned out before mid- afternoon

13.4 Components

cool after fires, portions of floors

fail and leave columns unbraced

14.2 Interior columns

stable for multi-story

lengths. Local floor

failure only.

13.5 Fire fails floor systems, impair

lateral column

support

14.3 Interior columns

stable for multi-story

lengths. Local floor

failure only.

14.4 Columns fail by

lateral forces from

thermal expansion

of floor system

section. Go to Vertical Progression

in Figure L-41

14.5 Temperature

gradient across column section fails cover plate welds or column splice. Go to Vertical Progression

in Figure L-41

14.6 Critical elevated temperature of steel components reached,

components fail to support loads. Go to Vertical Progression

in Figure L-41

.

Figure L-36. Collapse initiation scenarios.

L-37
Appendix L

o 14.1 Localized Failure at Interior Columns: If the interior columns failed just

after impact, this likely resulted in a local failure only, since the building continued to stand for almost 7 hours after WTC 1 collapsed. This failure could have progressed

vertically upward to the roof level within the bays immediately adjacent to the failed columns, yet from the northern vantage point of the photographic and videographic

observations, would not have been visible.

- 13.3 Interior Columns Remain Standing But Damaged: If interior columns were weakened by damage from debris, but retained sufficient capacity to carry their loads,

then additional loading and/or fire effects would have been required to cause their failure. Debris impact may have damaged the structural steel fireproofing without significantly

deforming the structural component.

11.2 Initiating Components Fail Due to Fire Effects: Fires had been burning in WTC 7 for many

hours, as observed in the photographic and videographic records (see section L.2). The initiating event may have been caused by fire effects on structural components.

• 12.3 Components on Floors With Burned Out Fires: If the initiating components failed

from fire effects, then locations where fires had burned out by mid afternoon could possibly

been affected by the cooling which occurs after a fire. No fire was observed or reported in the afternoon on Floors 1-5, 10, or above Floor 13.

- 13.4 Floor Systems Fail: The cooling that may have occurred as the fires burned out in an area may have generated thermal contraction forces, which may have induced tensile forces at floor-to-column connections.

o 14.2 Unbraced Columns: If floor systems failed, one or more columns may have

lost lateral bracing. At a floor where fires were noted, interior columns were

comprised of W14x730 cores and reinforcing plates, and could support several

stories unbraced without failure. As an example, the column capacity curve of

column 79 between Floors 5 to 9 is shown in Fig. L-37. Column load-carrying

capacities shown in this figure are based on the AISC column capacity formulas

(AISC 2001 ). The column is not very sensitive to the number of stories of unbraced

column length, K. This column, which had a service load stress of approximately

2 1 ksi, would be approaching its load carrying capacity for an unsupported length of

four stories if it was also subject to a unifonn temperature of 500 °C.

• 12.4 Components on Floors With Fire: If the initiating components failed because of fire

effects, then locations with uncontrolled fires would be more likely for the initiating event. From available data of fire locations in WTC 7, likely locations would include Floors 5, 6, 7,

8, 9, 11, 12, and 13. No fires were observed on Floor 5, but the lack of windows and the

presence of fuel systems on the south, west, and north floor areas indicate that fire should be

considered as a possibility on this floor.

- 13.5 Floor System Failure: The fires could have caused the failure of portions of one or more floor system and its framing connections.

L-38
Interim Report on WTC 7

T = 0

T - inn

T1 = ?nn OKJKJ

T = 700

Existing —B— K=1 —©— K=2 —e— K=4

0.5 1.5 2.5

Figure L-37. Column 79 capacity versus temperature and unbraced length K.

o 14.3 Unbraced Columns: If floor systems failed, one or more columns may have

lost lateral bracing. See 14.2 for discussion.

- 13.6 Columns, Transfer Girders or Transfer Trusses Fail: The fires could have

failed interior columns, transfer girders, transfer trusses, or their framing connections.

o 14.4 Lateral Displacements: Fire effects may have caused column instability failure by lateral displacements from asymmetric thermal expansion of the floor system. Such thermally-induced displacements must overcome the restraining effect of the remaining floor system against further lateral deflection of the column.

o 14.5 Temperature Gradients: Fire effects may have caused the failure of columns

and other components through the forces induced by temperature gradients through

their cross section. Bending and shear forces may be induced that are sufficient to yield either the column splice or reinforcing plate welds. Analysis of a one-story

segment of interior column 79 indicates that the cover plate weld would begin to yield at a mean temperature of 490 °C with a 200 °C gradient across the section, as shown in Fig. L-38. Other mean temperature and gradient combinations may also cause this type of failure.

L-39
Appendix L

400°C

Temperature Distribution across Column Section

0.5 in. Welds

26 in. X Sfn. Cover

Plate

Col. Lateral Deflection vs. Temperature

1 1 1

- Welds reach yield

'i I

,1'1

1

'1

li

ii —

'1 strength at 600 deg >

1

1 '1 200 400 600 Maximum temperature (deg C) Axial

thermal expansion

Creep

Total Displacement

Shear Flow

Location (ft) Weld Capacity Shear Flow Weld Capacity

Figure L-38. Effects of temperature gradient on interior column 79.

o 14.6 Uniform High Temperatures: If initiating event components were sufficiently

exposed to fire effects to be uniformly heated to elevated temperatures, the steel strength would be reduced below that required to support the load. Figure L-39

shows that for interior columns subject to service loads (shown as approximately

20 ksi of compressive stress), uniform steel temperatures of approximately 570 °C

would result in column failure.
Interim Report on WTC 7

Yield Strength vs. Temperature

100 200 300 400 500 600 700

25 T 725 Temperature (deg C)

Fy

Example of stress in steel

Figure L-39. Steel strength versus temperature.

L.3.3 Vertical Progression Scenarios

After the initiating component or components failed, there must have been a progression of the failure from the initiating event to other locations. To reflect the observed failure of the east penthouse, the

failure likely progressed vertically upwards. Figure L^O shows possible vertical progression scenarios.

The initiating component could have failed by any of the failure sequences listed under the collapse

initiation scenarios in Fig. L-36. This component could have been one of the columns under the east penthouse. It could also have been one of transfer trusses #1 or #2 under the east penthouse.

A collapse mechanism model was created to capture possible collapse initiation at the roof and the east penthouse. The model seeks to simulate only the kinematics of the collapse mechanism when columns

are removed. Several columns were tested for removal. The resulting geometry change was then

compared to the observed collapse of WTC 7.

VI. 1 Perimeter Columns Fail: Had the initiating component been any perimeter column, most likely it would have been at floor levels with debris impact damage (possible range extends from the ground level up to floors 15 to 20) or the floors possibly experiencing fire (Floors 5, 6, 7, 8, 9, 11, 12, or 13).

• V2.1 Collapse Does Not Progress: If a group of perimeter columns failed, the perimeter

framing above this area would have redistributed its loads, due to the redundancy of the moment frame.

V1.2 Core Columns Not Directly under East Penthouse Fail: Had the initiating component been a core column that was not under the east penthouse, most likely it would have been at floor levels with

debris impact damage (possible range extends from the ground level up to Floors 15 to 20) or the floors
Appendix L

V1 .1 Any perimeter

column fails V2.1 Collapse does

not progress vertically

Initiating event

scenarios

from 14.4 to 14.6

Vl .2 Core columns

not directly under east penthouse fail at floors 5-8 or 11-13

VI .3 Truss #1 or #2, east transfer girder, or columns 78 or 78A fail

VI.4 Interior columns

79, 80, and/or 81 fail at floors 5-13

V2.2 Collapse does

not progress vertically

V2.4 Collapse does

not progress vertically, need additional columns to fail - go to VI.

V2.5 Collapse progresses vertically at columns 76, 77, 78

V2.7 Collapse progresses vertically

V2.3 Collapse V3.1 Screenwall or progresses vertically west penthouse

collapse first

V3.2 Screenwall or west penthouse

collapse first

V3.3 East penthouse

collapses as seen in video records if columns 79, 80, or 81

also fail. Go to Horizontal Progression

in Figure L-41

V2.6 Collapse does

not progress vertically, need additional columns to fail - go to VI.

3

V3.4 Screenwall or west penthouse

collapse first

V3.5 East penthouse

collapses as seen in video records. Go to Horizontal Progression

in Figure L-41

Figure L-40. Vertical collapse progression scenarios.

possibly experiencing fire (Floors 5, 6, 7, 8, 9, 11, 12, or 13). However, a core column may have failed following the failure of adjacent columns or framing members.

• V2,2 Collapse Does not Progress: If core columns failed, the loads above the failed columns may have been redistributed to adjacent columns through the core floor system. If the loads could not be redistributed, then additional failures in one or more components

would have been necessary to progress the collapse.

• V2.3 Collapse Progresses: From this initial failure, the portion of the column above the

failure could have fallen, progressing the failure vertically upwards.

L^2
Interim Report on WTC 7

• V3.1 Something Else Besides East Penthouse Observed to Collapse First: Had the failure of core columns progressed upwards, then the first exterior sign of the internal failure likely

would have been seen in the screenwall or west penthouse, which are located above the core columns. A collapse mechanism analysis performed for the removal of columns 61, 64, 67,

70, and 73 produced geometry changes that differed from the observed collapse. For the

scenario where each of these columns fails and the failure progresses upwards to the roof line as the adjacent floors cannot redistribute the loads, the screenwall or the west penthouse

collapses, and no kink develops in the east penthouse (see Figs. L-41, h-A2, and L^3).

Image of Penthouse Kink Column 73 Removed

Figure L-41. Geometry changes for removal of column 73.

Column 70 Removed Column 67 Removed

Figure L-42. Geometry changes for removal of columns 70 and 67.

L^3
Appendix L

Column 64 Removed Column 61 Removed

Figure L-43. Geometry changes for removal of columns 64 and 61.

V1.3 Truss #1 or Truss #2 East Transfer Girder, or Columns 78 or 78A Fail: Had the initiating

component been truss #1 or truss #2, most likely there would have been debris impact damage or possibly fires at Floors 5 or 6. However, truss #2 failure could have followed the failure of the east

transfer girder or columns 78 or 78A.

• V2.4 Collapse Does Not Progress: If truss #1 or #2 failed, the floor framing, including the Floor 7 diaphragm, may have redistributed the loads to adjacent columns. Had this occurred,

additional failures in one or more components would have been necessary to progress the

collapse. For instance, the columns 76, 77, 78, 79, 80, or 81 may also have failed, and the combined effect of both component failures could have been sufficient to overcome the

supporting strength of the floor systems.

• V2.5 Collapse Progresses: If truss #1 failed, column 76 would lose its support at Floor 7, and the failure could have progressed vertically upwards if the floors could not redistribute

column 76 loads. If truss #2 failed, columns 77 and 78 would lose their support at Floor 7, and the failure could have progressed vertically upwards if the floors could not redistribute

the loads from columns 77 and 78.

o V3.2 East Penthouse Collapses Differently Than Observed: If truss #1 or truss #2

failed and the failure progressed vertically upward to the roof level, the exterior

deformations observed in the roof structures would be different from what was

actually observed. Column 76 supported the west side of the east penthouse and the

east end of the screenwall. A collapse mechanism analysis performed for the

removal of column 76 produced a geometry change that shows the west side of the

east penthouse and the east end of the screenwall deflecting downward

(see Fig. L-44).

L-44
Interim Report on WTC 7

Column 76 Removed

Figure L-44. Geometry changes for removal of column 76.

o V3.3 East Penthouse Collapses as Observed: Had the failure of columns 76 or 77

and 78 been followed by the failure of columns 79, 80, or 81, such that the failure of

column 79, 80, or 81 progressed upwards, while the vertical progression of failure above columns 76, 77, and 78 was arrested, then the first exterior sign of the internal

failures could have been observed at the center of the east penthouse roof

V1.4 Interior Columns 79, 80 or 81: Had the initiating component been column 79, 80 or 81 , most

likely the failure would have occurred at the floors possibly experiencing fire (Floors 5, 6, 7, 8, 9, 11, 12, or 13).

• V2.6 Collapse Does Not Progress: If only one of columns 79, 80, or 81 failed, the floor systems above the failure area may have redistributed the column loads to adjacent columns.

Had this occurred, additional failures in one or more components would have been necessary

to progress the collapse vertically upwards. For instance, both columns 79 and 80 may have

failed, and the loads of both columns could have been sufficient to overcome the supporting

strength of the floor systems.

• V2.7 Collapse Progresses: If only one of columns 79, 80, or 81 failed, the floor systems

above the failure area may have not been able to redistribute the column loads to adjacent

columns. The floor system above Floor 7 had beams and girders, concrete slabs on metal

deck, wire mesh in tenant floor areas, and rebar in the core area slabs. These floor systems do

not appear to have sufficient bending or catenary action to redistribute loads for failure of

column 79, 80, or 81. A calculation of the catenary action that might be developed by the beams and girders framing into column 79, assuming the floors try to redistribute the loads

above the area of column 79 failure, found that the girder connections reach their capacity at approximately 10 percent of the estimated service loads and the connections probably fail at approximately 25 percent of the service loads. If the floor-to-column connections had not

L-A5
Appendix L

failed, the beams would have started to yield axially at approximately 40 percent of the

service load present.

o V3.4 East Penthouse Collapses Differently than Observed: The collapse could

have progressed upwards, but the failure caused in the east penthouse could be

different than what was actually observed.

o V3.5 East Penthouse Collapses as Observed: Had the failure of the column

progressed upwards, then it could have been reflected in the observed collapse of the

east penthouse, which sits directly above columns 79, 80, and 8 1 . Also, the kink

observed in the roof of the east penthouse was in line with these columns. A collapse

mechanism analysis performed for the removal of column 79 produced a deformed

shape with a kink in the roof of the east penthouse (see Fig. L-45). This is a possible

collapse scenario.

Column 79 Removed

Figure L-45. Geometry change for removal of column 79.

L.3.4 Horizontal Progression Scenarios

After the east penthouse was observed to sink into the building core, approximately five seconds lapsed

before the screenwall and west penthouse were observed to also sink into the building core. The

screenwall and west penthouse movements occurred almost simultaneously with the global collapse of the

structure. From these external observations, it appears that after the vertical progression failure on the

east side of the building, the failure progressed horizontally across the core. The horizontal progression

of the collapse could have started due to any of the likely vertical progression scenarios, which are shown

in Fig. L-A6.

L~46
Interim Report on WTC 7

Debris from vertical collapse impacts and

damages or severs

adjacent

components

H1.1 Any typical floor system above floor 7 fails

H1.2 Floors 5 and/or 7 fail and exert lateral forces on adjacent columns

H1.3Truss#1 fails, column 76 fails, floors 5/7 exert lateral forces on

adjacent columns

HI.4 Truss #2 and/or

east transfer girder fail, columns 77 and 78 fail, floors 5/7 exert lateral forces on adjacent

columns

H2.1 Horizontal progression

unlikely

H2.2 Structure can resist horizontal loads and

progression stops, need

additional failures, go to HI .3

H2.3 Lateral forces cause many core columns to fail; leads to sudden onset of global instability

H2.4 Collapse does not progress horizontally

H2.5 Collapse progresses

horizontally at columns 73, 70,

67, 64, 61; leads to sudden

onset of global instability

H2.6 Collapse does not progress horizontally

H2.7 Collapse progresses

horizontally at columns 74, 71

,

68, 65, 62; leads to sudden

onset of global instability

Figure L-46. Horizontal collapse progression scenarios.

The likely region in which the horizontal progression occurred is in the lower portion of the building,

around Floors 5 and 7. Floor 5 had a 14 in. reinforced concrete slab on metal deck. The slab was heavily

reinforced, and contained steel WT sections embedded in the slab. The WT sections were arranged in a diagonal pattern, like a horizontal truss, within the plane of the floor between the perimeter and core columns. Floor 7 had an 8 in. reinforced concrete slab on metal deck with rebar in each direction. The

beams between interior columns at Floors 5 and 7 were much larger than at other floors, and the beam-tocolumn connections were able to transfer more of the beam axial and bending load capacity. These strong

lateral ties between the interior columns may have been able to impose lateral displacements on adjacent

columns. Transfer trusses and girders between Floors 5 and 7 transferred loads from the columns above

Floor 7 to the foundation columns below Floor 5.

Assuming that a vertical collapse of one or more bays occurred over the height of the building, a large

pile of debris would have fallen on Floor 7 and below. Such a large amount of debris is likely to have
Appendix L

severed the Floor 7 slab and damaged or severed any transfer tmss or girders in the vicinity. For a

vertical collapse on the east side of the building, transfer trusses #1 and #2 and the east transfer girder

may have been damaged, particularly the east diagonals of the trusses. The scenarios below describe

possible responses of Floors 5 and 7 following a vertical collapse of one or more bays.

Hl.l Floor Systems above Floor 7: Typical tenant floors above Floor 7 were constructed with concrete

slabs metal deck with wire mesh reinforcement. The steel framing coimections were designed for shear

loads only, though they could likely resist some degree of tensile catenary forces.

• H2.1 Collapse Does Not Progress: For any interior column failure above Floor 7, the tenant and core floor systems are not able to develop sufficient axial tensile loads for imposing lateral deflections on adjacent columns. It is likely that the floor system within a bay will fail before a column failure is propagated horizontally to adjacent columns.

H1.2 Floors 5 and 7: Floors 5 and 7 were thicker and more heavily reinforced than the typical floor systems, and may have been subjected to a large debris load from a vertical collapse within one or more

bays.

• H2,2 Collapse Does Not Progress: Floors 5 and 7 may fail at connections to adjacent

columns before developing any tensile forces large enough to cause other column failures through lateral displacements, halting the horizontal progression.

• H2,3 Collapse Progresses: Floors 5 and 7 may impose large tensile forces at the adjacent

columns to cause lateral displacements that fail the columns. The failure mechanism could occur at the column splice, located just above Floor 5 and Floor 7, rather than through the column section. The simultaneous occurrence of column instability in many core columns

would cause a sudden and large change in the structural system capacity.

H1.3 Truss #1: If one of the diagonals of truss # 1 (see Fig. L-47) was damaged or severed by collapse

debris from the vertical progression, there would be a horizontal force developed in the Floor 7 slab as column 76 became unstable. The floor beam between column 76 and column 73 would try to restrain column 76 movement through tensile forces to column 73.

• H2.4 Collapse Does Not Progress: The horizontal tensile force would tend to pull the line of columns 73, 70, 67, 64, and 61 towards the east. The continuity of the Floor 7 slab and the presence of braced frames around the north core column line makes the simultaneous lateral displacement of the core columns less likely, as such displacements within a rigid slab may

similarly displace other columns, including perimeter columns.

• H2.5 Collapse Progresses: The failure of column 76 may create its own vertical collapse,

due to the inability of the floor systems above to redistribute the loads and fail at the column

splices near Floors 5 and 7 as shown in Fig. L-48. If column 76 cannot be restrained and

there is a vertical collapse of the surrounding bay, it would cause a debris pile at the lower

floors which may then destabilize adjacent columns.

L^8
Interim Report on WTC 7

Splices Fail

in Bendina

Truss #1

XZ3

Column Fails

in Bendina 0

Figure L-48. Horizontal progression mechanism for truss #1 failure.

L-49
Appendix L

H1.4 Truss #2 and/or East Transfer Girder: If one of the diagonals of truss # 2 and/or the east transfer girder was damaged or severed by collapse debris from the vertical progression, there would be a

horizontal force developed in the Floor 7 slab as columns 77 and 78 became unstable.

• H2.6 Collapse Does Not Progress: The Floor 7 slab may fail at adjacent columns prior to imposing lateral displacements sufficient to fail the columns or their splices.

• H2.7 Collapse Progresses: The horizontal tensile force would tend to pull the line of

columns 74, 71, 68, 65, and 62 towards the east. The general absence of the Floor 7 slab and

braced frames around the center core column line, due to the presence of elevators shafts, creates a more likely scenario for the simultaneous lateral displacement of the center core columns without similarly displacing other core columns. The possible result is a failure of

all the columns at their splices, as shown in Fig. L^9.

L.3.5

0 <^

^ 0 0

Sglices Fail in Bending

Column Fails in Bending

Figure L-49. Horizontal progression mechanism for truss #2 failure.

Summary of Working Collapse Hypothesis

G

The working collapse hypothesis has been developed around four phases of the collapse that were

observed in photographic and videographic records: the initiating event, a vertical progression at the east side of the building, and a horizontal progression from the east to west side of the building, leading to global collapse.

From an analysis of the observed collapse sequence, the following general sequence of events appears

possible:

1. Debris damaged the south face of the perimeter moment frame and some interior core framing on the south side. The debris impact severed approximately a quarter to a third of

L-50
Interim Report on WTC 7

the south face perimeter columns. The damaged floors are less certain, but reports indicate

they occurred between the ground and up to Floors 15 or 20. The extent of damage, both

structural and to fireproofmg, of core framing is not known, but damage to elevator cars and

shafts was reported to have occun-ed around columns 69 to 78 at Floors 8 or 9.

2. Fires were observed after the collapse of WTC 1 . Fires were observed after 2 pm on

Floors 7, 8, 9, 11, 12, and 13. Fires were not observed on Floor 5, but this may be due to the

lack of windows. The presence of a fuel distribution system and the possibility of damage at the south face fi'om WTC 1 debris impact, indicates that fires may have been present on Floor

5.

3. The initiating event may have included a number of structural components, though the

relative role of impact damage and fire need further investigation. Possible components that may have led to the failure of columns 79, 80, and/or 81 include interior columns 69, 72, 75,

78, and 78A, the east transfer girder (which supports column 78A and frames into transfer

truss #2), and adjacent framing and floor systems.

4. A vertical collapse appears to have occurred after interior columns 79, 80, and/or 81 failed. This failure mechanism would progress vertically upward within the failed bay to the roof

level, as analysis indicates that the floors would not be able to redistribute their loads.

5. The debris from a 40-story vertical collapse on the east side of the building would fall down

onto the strong diaphragms at Floors 5 and 7 and possibly onto transfer trusses #1 and #2,

and/or the east transfer girder. Damage and loading on these floors and transfer components

would generate lateral forces which would cause the failure of the remaining core columns.

The horizontal progression requires further analysis and investigation, but observations

indicate that the remaining core columns appeared to fail almost simultaneously,

approximately 5 second after the east penthouse failed.

6. The core columns failed and redistributed loads until the building loads could no longer be

supported. Once the core columns failed, the cantilever girders which supported the north

facade also failed. The remaining perimeter columns at the east, south, and west facades were either left unsupported or were pulled down with the interior collapse. The global

collapse occurred with few external signs prior to the system failure.

The working hypothesis, for the collapse of the 47-story WTC 7, if it holds up upon ftirther analysis,

would suggest that it was a classic progressive collapse that included:

• An initial local failure due to fire and/or debris induced structural damage of a critical column, which supported a large span floor area of about 2,000 ft", at the lower floors (below Floor 14) of the building,

• Vertical progression of the initial local failure up to the east penthouse bringing down the

interior structure under the east penthouse, and

• Horizontal progression of the failure across the lower floors (in the region of Floors 5 and 7

that were much thicker and more heavily reinforced than the rest of the floors), triggered by

L-51
Appendix L

damage due to the vertical failure, resulting in disproportionate collapse of the entire

structure.

The working hypothesis will be revised and updated as results of ongoing, more comprehensive analyses

become available.

L.3.6 Technical Approach for Analysis of the Working Collapse Hypothesis

There are many possible collapse scenarios that have been postulated in the preceding section. Many of

the scenarios will not produce the observed sequence of global collapse events and can be classified as unlikely. Likely collapse scenarios will be identified through analyses that test the postulated phases of

collapse against observations. It is equally important to test scenarios that are not predicted to match the observed data. The testing of the postulated collapse scenarios will be conducted through hand

calculations, simplified nonlinear thermal-structural analysis, and full nonlinear thermal analysis.

L.4 REFERENCES

American Institute of Steel Construction Inc., Manual of Steel Construction, Load and Resistance Factor Design, Third Edition, 2001. Chapter 16-E.

McAllister, T., ed. 2002. World Trade Center Building Performance Study: Data Collection, Preliminary

Observations, and Recommendations. FEMA 403. Federal Emergency Management Agency.

Washington, DC, May.

Salvarinas, John J. 1986. Seven World Trade Center, New York, Fabrication and Construction Aspects,

Proceedings ofthe 1986 Canadian Structural Engineering Conference, Vancouver, Canadian Steel Construction Council, Ontario. February 24-25.

L-52
Table of Contents

List of Figures M-ii

Appendix M

Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire M-1

M.l Summary M-1

M.2 Introduction and Review M-1

M.3 Structural Model M-1

M.4 Material Properties M-2

M.5 Loading M-4

M.6 Results M-6

M.7 Conclusions M-11

M.8 Note M-12

M.9 References M-12

M-i
List of Figures

Figure M-1 . Yield strength of steels used in model M-3

Figure M-2. Modulus of elasticity of steels used in model M-3

Figure M-3. Mechanical properties of concrete slab as modeled M^

Figure M^. Temperature distribution of "real fire" scenario 2-04 M-5

Figure M 5. Deformed shape at 1,273 K M-7

Figure M-6. Horizontal force between column and floors versus temperature M-7

Figure M-7. Vertical deflections of floors versus temperature M-8

Figure M-8. Overall deflected shape for fire scenario 2-04 M-8

Figure M-9a. Maximum deflection of floors 95-98 for fire scenario 2-04 M-9

Figure M-9b. Maximum deflection of floors 95-98 for fire scenario 2-04: details of

Fig. M-9a at high temperatures M-9

Figure M-10. Perimeter column lateral deflection for fire scenario 2-04 M-10

Figure M-1 1 . Horizontal force at connection between floor and internal column for

fire scenario 2-04 M-10

Figure M-1 2. Horizontal force at connection between floor and external column for

fire scenario 2-04 M-11

Figure M-13. Overall deflection of 12-floor model (floors 91-102) subjected to gravity loads

and with floors 95 and 96 under conventional fire M-13

Figure M-14. Details of floors 93-98 for the 12-floor model shown in Fig. M 13 M-14

Figure M-1 5. Details of temperature distribution of floors 92-95 for 12-floor model M-14

M-ii
Appendix M

Interim Report on 2-D Analysis of the WTC Towers Under

Gravity Load and Fire

M.1 summary

A two-dimensional (2-D) finite element model is developed to provide insight and evaluate some aspects

of a possible collapse sequence for the World Trade Center (WTC) towers. For a prescribed temperature

distribution that corresponds to a two-story, quarter-span fire, and for a three-story fire derived from fire dynamics simulation, diagonals of the heated trusses buckle inelastically, causing considerable sag in the

fire floors. This behavior puts a high-tension demand on the truss connections to the perimeter column,

which remains at moderate temperatures in this model and does not experience buckling. Because neither

the prescribed nor the derived temperature distributions are necessarily representative of the actual fire, and the material properties are approximate, further work is needed to evaluate the collapse sequence and

develop findings regarding the actual event.

Within days of the collapse of the WTC towers on September 11, 2001, publications postulating the mechanism of the collapse began to circulate. A substantial effort was launched by the Federal

Emergency Management Agency (FEMA) and the American Society of Civil Engineers (ASCE),

culminating in a preliminary building performance study (McAllister 2002). Quintiere et al. calculated

the elastic buckling strength of a single diagonal of a floor truss, assuming pinned end conditions, and

suggested that the buckling of such thin members exposed to fire might have initiated the collapse

(Quintiere et al. 2002). More recently, Usmani et al. performed a series of 2-D, nonlinear finite element

analyses of a 12-story vertical frame that comprises a perimeter column and, at each floor, a truss and

floor slab supported by the column and the tower core (Usmani et al. 2003). The temperature distribution

in the steel and concrete members was characterized by an assumed time-dependent profile. Usmani et al. concluded that column instability caused by the loss of bracing normally provided by floors led to overall

structural collapse (Usmani et al. 2003).

The objective of this report is to present a simplified analysis approach to evaluate some aspects of the

collapse sequence of the WTC towers. The analysis is based on a 2-D model that is simple and can be

easily used to evaluate a wide variety of conditions. The structural system was modeled independently of

connection details. At this stage, cormections are the object of a separate analysis that can draw on the

results presented here concerning demand upon connecfions at various stages of fire development.

The vertical plane considered in the model includes perimeter column 109 on the North face of WTC 1, and five longitudinal floor trusses and slabs (floors 94 to 98). The center of the airplane impact was at

M.2 INTRODUCTION AND REVIEW

M.3 STRUCTURAL MODEL

M-1
Appendix M

floor 96, and column 109 was the intact column closest to the edge of the initial damage zone

(McAllister 2002). The column extends 22 m (72 ft) to a height of six floors, and both its upper and

lower ends are pinned, with the upper end free to translate vertically. The upper chords of the floor

trusses are simply supported at the internal end, and connected to the perimeter column by hinges. In the

actual structure, a double floor truss carries a tributary floor slab 2 m (80 in.) wide and is supported by

two perimeter columns, whereas in the present model a single truss supported by a single perimeter

column carries a 1 m (40 in.) wide slab. The floor spans 18.3 m (60 ft) from the perimeter column to the

core. The model is similar to that of Usmani et al. (2003), except it has fewer floors.

The principal reason for including only five floors in the analysis is to have the simplest model that will

still capture salient features of the collapse of the towers. The fire applied to the model only heats two

floors, and the remaining floors remain cool and provide lateral restraint to the perimeter column under

study. Since the ends of the perimeter column in this simple model are hinged, the model ignores the

rotational restraint supplied by the continuous column if additional floors are considered. Thus the short model is less stiff than a taller model (such as the I2-floor model developed by Usmani [2003] and would

buckle sooner (or at a lower mechanical or thermal load), if global buckling should occur at all. As far as

translational restraint is concerned, only a very small amount of lateral bracing can have a tremendous

effect on the buckling strength (Winter 1958), and the cool floors one or two stories away from the fire can be replaced by a support that does not allow horizontal translation. One additional reason for including only five floors is to provide a guide for and allow comparisons with the results of a three- dimensional study, where several full floors are included. The size of the three-dimension model is a concern.

The trusses, slabs and the column that supports them are simulated by three-node beam finite elements,

capable of modeling a wide variety of cross sections, with a mesh density and number of integration

points specified by the user. One particularly attractive feature of these elements is the capability of

supporting linear temperature gradients across their section and along their length.

M.4 MATERIAL PROPERTIES

The various steels range in nominal yield strength from 250 MPa (36 ksi) in the floor trusses to 450 MPa

(65 ksi) in the column (McAllister 2002). They are all modeled by bilinear stress-strain curves, with a tangent modulus about 0.5 percent of the elastic modulus. Figures M-1 and M-2 show the steel properties for the temperature range used in the analysis. Usmani et al. (2003) used similar steel properties.

The lightweight concrete slab is also modeled as a bilinear, ductile material (Fig. M-3), with compressive

strength of 20 MPa or 3,000 psi (McAllister 2002). The top chord of the floor truss is assumed to act in a perfectly composite way with the slab and allow the tensile strength at the bottom of the slab to be equal

in magnitude to the compressive strength at the top. This choice of a simple, bilinear material

overestimates the tensile capacity of the slab. As well, the simplification inherent in transforming the

steel top chord into an equivalent concrete section disregards the differential thermal expansion between

steel and concrete. A more accurate concrete model (currently being developed) may show slab failure or

a smaller horizontal tension at the connection between floor and column than the present results.

M-2
Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire
Appendix M

Compressive strength, MPa

Modulus of elasticity, GPa

M.5

273 373 473 573 673 773 873 973 1073 1173 1273

Temperature (K) Figure M-3. Mechanical properties of concrete slab as modeled.

LOADING

The floor slabs are acted upon by a dead load of 3.3 kPa (70 psf) and a live load of 720 Pa (15 psf). The

column load, determined by a linear, static finite element analysis of the global, damaged structure, includes the weight of the floors above and a surcharge due to load transfer from the columns damaged or missing after the airplane impact (Appendix D, Section D.2.4 of this report). The top of the column is loaded by a 1,100 kN (250 kip) axial compressive force and a 2,000 N-m (18 kip-in.) clockwise moment

(compared to 540 kN or 120 kip, and 1,800 N-m or 16 kip-in. counter clockwise moment before damage).

In addition, the column self-weight is applied along its length. For comparison, Usmani et al. (2003) used

loading consistent with the FEMA report (McAllister 2002) and applied 40 percent of the gravity loads of

the tributary floor strips above the model to the top of the perimeter column.

The behavior of the structure and its eventual collapse are greatly influenced by thermal loads. This

report first performs an analysis based on a conventional fire, which provides a useful first approximation

to the behavior of the building in fire. For comparison with the work of Usmani et al. (2003), a single

temperature distribution T represented by an exponential function of time t, with a reference temperature

To = 300 K, is used. The time rate of change of the temperature, represented by coefficient a = 0.005,

depends, among other factors, on the location and intensity of the fire, and the quality of the insulation.

m^T,+(T^^^-T,)(l-e-'") (1)
Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire

A two-floor Are, with maximum temperamre T^ax = 1,273 K, heats the structure on floors 95 and 96, over

the quarter-span closest to the perimeter column. Over that span, the slab of floor 95 is uniformly heated,

whereas the slabs of floors 94 and 96 have linear temperature gradients across their thickness, with the bottom of slab 94 and the top of slab 96 remaining at 300 K at all times. In the three-quarters of the span

not directly under fire, the temperature decreases linearly from the maximum at quarter-span to room

temperature at the core. Between floors 94 and 96, the column temperature is also described by Eq. (1), with Zmax = 400 K, whereas the rest of the column remains at 300 K at all times.

The finite element model was used further by applying to it a second temperature distribution (Fig. M^)

that corresponds to a more realistic, physics-based fire, generated by NIST's Fire Dynamics Simulator in a manner consistent with initial conditions appropriate for the WTC towers following the aircraft impacts.

The gas temperatures associated with the fire were used to calculate structural member temperatures and

temperature gradients, assuming an insulation thickness of 19 mm (3/4 in.) for truss members and 36 mm

(1.4 in.) for the perimeter column. The fire considered in this application is more widely spread than the conventional fire, and covers four floors, with the entire floor span heated. Floor 94 (lowest) remains

unheated, and the column is only moderately heated. Linear temperature gradients are modeled across the column section and the slab thickness of heated floors. The peak temperature of 1,230 K is obtained at the end of 25 temperature load steps, each 200 s apart.

ELEMENTS

U

P.OT FH

CPNFOR

miuii

P.FOP. PJIOH TKHPEPATURES

THIN=301. 674

THAX=i229

AN

\ A .

KAA.AAA,/yAA,AAA/xAA

J ^

MAR 18 2004

17: 03: 33

301.674 507.668 713.662 919.656 1126

404.671 610.665 816.659 1023 1229

Real temp2 BISO concrete slab Voce diag. birth & death

Figure M-4 Temperature distribution (K) of "real fire" scenario 2-04.

M-5
Appendix M

M.6 RESULTS

Nonlinear, static, large deformation analysis accounting for the magnification of flexural deflections due

to axial load (P-delta effect) was performed. Member stiffness matrices are updated during the analysis to account for the P-delta effect, and when a member stiffness gets close to zero, excessive lateral deflection occurs and the member is considered to have buckled. The first analysis proceeded in eight load steps, the first corresponding to gravity loads at the start of the fire (normal room temperature). Subsequent

steps occurred at 200 s intervals, with the maximum temperature attained, to within 1 K, at 1,400 s. As

required in the computation, the load steps are further divided into substeps (up to several hundreds).

Results for the conventional fire are shown in Figs. M-5, M-6, and M-7.

At room temperature, even under the severe load redistribution due to the damage caused by the airplane

impact, the structure still behaves linearly. The maximum floor sag is 35 mm (1.4 in.), causing the horizontal span to decrease and the column to pull in slightly. Approaching 200 s and a temperature of

9 1 5 K (the temperatures referred to in these results are the hottest temperatures in the structure at any

given time), the heated truss begins to show distress, especially in the compressed diagonal and vertical web members, which buckle inelastically. This means these heated steel members do not buckle

elastically, but rather reach yielding in compression. Buckling is then governed by the tangent modulus

of steel, which is about 0.5 percent of the elastic modulus, and the members immediately buckle after yielding, in the inelastic range. At 200 s, the maximum floor sag increases to 335 mm (13.2 in.), and the column is pushed out (peak of 38 mm or 1 .5 in.) by the thermal expansion of floor 95. At that time, the connection of floor 95 to the perimeter column experiences its maximum compression of 125 kN (28 kip).

Because slab 96 has a thermal gradient with its top surface at room temperature, its lateral expansion is much smaller than for slab 95, and its sag is larger. The connection between slab 96 and the column is always in tension (Fig. M-6). As expected, slab 94, heated at the top and cool at the bottom, bows

upward. As the temperature continues to rise, more of truss 95 web members buckle inelastically, and the

increasing sag begins to pull the column in. The horizontal deflection of the column becomes positive

(inward), and the connection force between column and floor 95 turns to tension. This inward movement

of the column relieves the tension in the connection between the column and floor 96. Further

temperature rise causes further weakening in truss 96, which eventually becomes active in pulling the

column in. At the peak temperature of 1,273 K, the maximum lateral deflection in the column (183 mm

or 3.3 in.) occurs at floor 96, inward, and the connection between the column and floor 96 experiences a tension of 185 kN (42 kip).

Under the second fire scenario, the structure exhibits similar behavior. Figures M-8, M-9, and M 10

show the resulting deflections. Inelastic buckling of the diagonals causes considerable vertical deflection

of the heated floors beyond a maximum temperature of 900 K. At 1,220 K, the sag of floors 96 and 97 overcomes the outward push on the perimeter column due to thermal expansion of floors, and pulls the

column inward. This transition causes the column to temporarily straighten up, causing the overall floor deflection to be less. Figures M-1 1 and M-12 show severe horizontal tension greater than 120 kN

(27 kip) at the connection of floors with the internal column (floor 95) and external column (floor 96).

M-6
Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire

Floor 98 r^KAAAA/VV\yV/\AAA-A7\A/

Floor 97

Floor 96

, Floor 95

\-^h/\/\/\/\/\/\/\/\/\/\/\/\/_\i;^/f<j

Floor 94

X

Figure M-5. Deformed shape at 1,273 K (not to scale).

200

165

130

95

a> 60 ut 25

re

I -45

-80

-115

-150

1 — Floor 95 — Flo<ar 96

_

i —

1

1

'

1

300 500 700 900 1100

Temperature (°K)

Figure M-6. Horizontal force (kN) between column and floors versus temperature (K).

M-7
Appendix M

300 500 700 900 1100 1300

Temperature (°K)

Figure M-7. Vertical deflections (mm) of floors versus temperature (K).

Figure M-8. Overall deflected shape for fire scenario 2-04 (not to scale).

M-8
Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire

0

-65

130

195

260

325

390

455

•520

585

650

-

i

_ —^^Vl

Wi —Floor 95 —FInor Qfi —^ Floor 97

loor

300 500 700 900

Temperature ("K)

1100 1300

Figure M-9a. Maximum deflection of floors 95-98 for fire scenario 2-04.

Figure M-9b. Maximum deflection of floors 95-98 for fire scenario 2-04: details of

Fig. M-9a at high temperatures.

M-9
Appendix M

E

E

u

o

a

+-> c

o

_N

oX

c

E

O

o

15

12

9

6

3

0

-3

-6

-9

12

15

— Fl<Dor 9

Dor 9

3or 9

4

—r- Fl<

Fli 5

D — Fl<DOT 9

Dor 9

7 — Fl« 8

750 850 950 1050

Temperature (°K)

1150 1250

Figure M-10. Perimeter column lateral deflection for fire scenario 2-04.

u

o

u.

15

£

o

N

750 850 950 1050 1150 1250

Temperature (°K)

Figure M-11. Horizontal force at connection between floor and

internal column for fire scenario 2-04.

M-IO
Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire

O

(0O

N

150

120

90

60

30

0

-30

-60

-90

120

150

— Floor 9A[ Floor 95

Floor 96

Floor 97 r — noor yo'

1

!

I

750 850 950 1050

Temperature ("K)

1150 1250

Figure M-12. Horizontal force at connection between floor and

external column for fire scenario 2-04.

NIST analysis of connections is ongoing and will indicate whether this or other connections can supply

the calculated demand, and if not, at what temperatures connection failures will occur. In this regard, the

loss of composite behavior of the concrete slab with the steel truss may occur at a temperature and strain level yet to be determined.

For comparison with Quintiere et al. ( 2002), the present results show that the truss diagonals buckle

inelastically, and there is considerable reserve strength after the first diagonal buckles. At the highest

temperatures analyzed, seven diagonal and the vertical web members had buckled in each of the floors heated by the conventional fire. This conclusion assumes that the various structural connections maintain

their integrity throughout the fire.

M.7 CONCLUSIONS

A model has been developed to provide insight and evaluate some aspects of a possible collapse sequence

of the WTC towers. Its results are subject to the following qualifications: (1) the approximate nature of

the material properties used, especially the concrete slab; (2) connection failures are not considered,

although information is provided on demand experienced by the connections; and (3) the model is twodimensional. In one of the two cases covered by this report, the temperature distribution of the members

is selected from among those assumed by Usmani et al. (2003). In the second case, the temperature

distribution is physically based, and was obtained by using the NIST Fire Dynamics Simulator with

reasonable initial conditions associated with a damaged tower. In both cases, the diagonals buckle

inelastically, causing considerable sag in the fire floors. This behavior puts a high-tension demand on the

M-U
Appendix M

column, which remains at moderate temperatures in this model (same temperature as in Usmani et al. [2003]), and does not experience buckling. This is the major difference between these results and Usmani

et al., even though the heated trusses in the present model are exposed to a much higher temperature and

the column to a more severe load that reflects load redistribution in the damaged structure. One possible

explanation for the difference is that failure modes may be sensitive to material properties.

M.8 NOTE

For confirmation, a 12-floor model (from floors 91 to 102) was developed and loaded with the same floor load and conventional temperature distribution described by Eq. 1 (Usmani et al. 2003) and mentioned

above. Compared to case 1 reported earlier, the hinged column end conditions, the heated floors (95 and

96) and the bending moment applied on top of the perimeter column are the same, but the axial

compression is reduced (950 kN or 210 kip) because of the fewer floors above the model. Results of the

12-floor model, shown in Figs. M-13, M-14, and M-15, are very similar to those of the 5-floor model,

thus confirming the discussion in Section M.3.

M.9 REFERENCES

McAllister, T., ed. 2002. World Trade Center Building Performance Study: Data Collection,

Preliminary' Obsen'ations, and Recommendations. FEMA403. Federal Emergency Management

Agency. Washington, DC, May.

Quintiere, J. G., M. DiMarzo, and R. Becker. 2002. A suggested cause of the fire-induced collapse of the

World Trade Towers, Fire Safety Journal, vol. 37, no. 7. October, pp. 707-716.

Usmani, A. S., Y. C. Chung, and J. L. Torero. 2003. How did the WTC towers collapse: a new theory,

Fire Safety Journal, vol. 38, no. 6. October, pp. 501-533.

Winter, G. 1 958. Lateral bracing of columns and beams, ASCE Journal ofStructural Division,

Proc. 1561. March. Also in ASCE Transactions 3044.

M-12
Interim Report on 2-D Analysis of the WTC Towers Under Gravity Load and Fire

DISPLACEHEHT

STEP=8

SUB =1

TIHE=1Z72

DNX =42.444

X

Usmarii temp BI30 composite slati Voce diag.l2 floors

Figure M-13. Overall deflection (not to scale) of 12-floor model (floors 91-102)

subjected to gravity loads and with floors 95 and 96 under conventional fire: at the maximum temperature of 1,272 K, the maximum deflection is 1.08 m.

M-13
1 DISPLACEMENT

STEP=8

SUB =1 TIHE=1272

'^K/V\/\y\/\/\/\y\/\AAA/\/\/\AN

HAY 28 2004 19:28:22

DHX =42. 444 -^l\/V\/\AAAyV\AAAAA/\A/

^KyV\yV\/V\AAAAAA/\A^^

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K AA/A A//\ /A AvA AA A A y\ /\ y\ /

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Usmani temp BISO co: aposite slab Voce diag.l2 floors

Figure M-14. Details of floors 93-98 for the 12-floor model shown

in Fig. M-13 (compare with Fig. M-5).

Figure M-15. Details of temperature distribution of floors 92-95 for

12-floor model: note temperature gradients across slab thickness of

floors 93 and 95, and cool column.

M-14