NIST Technical Note 1771

A Study of Thermal Behavior of a 

Composite Floor System in 

Standard Fire Resistance Tests

Dilip Banerjee

http://dx.doi.org/10.6028/NIST.TN.1771
2

NIST Technical Note 1771

A Study of Thermal Behavior of a 

Composite Floor System in 

Standard Fire Resistance Tests

Dilip Banerjee

Materials and Structural Systems Division

Engineering Laboratory

http://dx.doi.org/10.6028/NIST.TN.1771

October 2012

U.S. Department of Commerce 

Rebecca Blank, Acting Secretary

National Institute of Standards and Technology 

Patrick D. Gallagher, Under Secretary of Commerce for Standards and Technology and Director 
3

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

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

Such identification is not intended to imply recommendation or endorsement by the

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

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

National Institute of Standards and Technology Technical Note 1771

Natl. Inst. Stand. Technol. Tech. Note 1771, 46 pages (October 2012) 

http://dx.doi.org/10.6028/NIST.TN.1771

CODEN: NTNOEF
4

Table of Contents

Abstract............................................................................................................................... 5

1 Introduction...................................................................................................................... 6

2 Description of the Fire Resistance Tests.......................................................................... 6

2.1 Test Specimen........................................................................................................... 7

2.2 Furnace...................................................................................................................... 8

2.3 Instrumentation ......................................................................................................... 9

2.4 Test Procedure ........................................................................................................ 10

3 Experimental Results..................................................................................................... 11

3.1 General Observations of Tests................................................................................ 11

3.2 Furnace Thermal Environment ............................................................................... 11

3.3 Unexposed Concrete Slab Surface Temperatures................................................... 13

3.4 Steel Temperatures.................................................................................................. 14

3.5 Variability in Experimentally Recorded Temperatures.......................................... 16

4 Heat Transfer Modeling................................................................................................. 20

4.1 Methodology........................................................................................................... 20

4.2 Thermophysical Properties, Boundary and Initial Conditions................................ 22

4.3 Interpolation of Temperatures................................................................................. 23

5 Evaluation of the Heat Transfer Model ......................................................................... 24

5.1 Test 1....................................................................................................................... 24

5.2 Test 2....................................................................................................................... 27

5.3 Effect of SFRM Dislodgement ............................................................................... 32

6 Summary........................................................................................................................ 34

7 References...................................................................................................................... 36

8 Acknowledgment ........................................................................................................... 37

Appendix A: Thermophysical properties.......................................................................... 38

1. Steel........................................................................................................................... 38

2. Concrete .................................................................................................................... 40

3. SFRM........................................................................................................................ 41

Appendix B: Finite Element Model of the Composite Floor System............................... 43

Appendix C: Log of Observations in Test 1 ..................................................................... 45

Appendix D: Log of Observations in Test 2..................................................................... 46
5

Abstract

The National Institute of Standards and Technology (NIST) conducted a series of four 

standard fire resistance tests of the composite floor system used in the World Trade 

Center (WTC) towers. These tests were conducted as part of NIST’s investigation into 

the collapse of the WTC towers. The main objectives of these tests were to understand 

the effects of three key factors on the fire behavior of composite floor systems during fire 

exposure. These factors were: a) thickness of fireproofing on steel, b) test restraint 

conditions, and c) scale of the test (e.g., full scale vs. reduced scale). Two tests were 

conducted using full-scale test specimens while two other tests were conducted using 

reduced-scale test specimens. In this report, only the results from the two full-scale tests 

are discussed and compared against numerical predictions. These two thermally-similar 

fire resistance tests were conducted by the Underwriters Laboratories (UL) of Canada by 

exposing a composite floor system to the heating profile (time-temperature curve) 

specified in the ASTM E119 standard. In these tests, the fire environment was well 

controlled and measured. Also, the temperature and deflection data were recorded at 

sufficient number of locations to characterize the structural performance of the test 

assemblies over time. Since a large set of well controlled measurements (both thermal 

and structural) are available, data from these tests were used for development and 

evaluation of numerical models to understand the thermal and structural behavior of 

composite floor systems during fire exposure.

This report summarizes the characterization of the furnace thermal environment and 

thermal behavior of members during the fire tests. Additionally, numerical results from a 

heat transfer finite element model of a representative section of the floor system are 

compared against data obtained in the full-scale tests to evaluate the finite element heat 

transfer model. Future work will focus on determining uncertainties in predicting 

temperatures of structural members using experimental data collected during NIST’s 

WTC fire resistance tests.
6

1 Introduction

Predicting structural behavior under fire requires an accurate modeling of (a) thermal 

environment in a fire, (b) heating of structural elements exposed to hot gases from fire, 

and (c) structural response to both thermal and gravity loads. There are uncertainties 

associated with the computed results in each of these steps, which propagate from one 

step to the next. In analyzing uncertainty propagation, the goal is to determine how

random variation, lack of knowledge, or error affects the sensitivity, performance, or 

reliability of the system that is being modeled. The predicted results need to provide 

time-dependent temperatures in structural members during a fire event, along with the 

associated uncertainties. 

A comparison of numerical predictions of the temperature field against carefully 

controlled fire resistance experiments can provide insight into the uncertainties associated 

with prediction of the temperature histories of structural elements exposed to fire. 

However, data from such well controlled full-scale fire resistance tests are difficult to 

obtain from the literature. Since a rich set of data from the National Institute of Standards 

and Technology (NIST)’s fire resistance tests of composite floor systems is available, 

validation efforts have been focused on these tests. Toward this end, a finite element heat 

transfer model of the composite floor truss section was constructed in ANSYS [1] to 

model the heating profile of the structural members exposed to the standard fire. 

Although the details of each fire resistance test are described in ref. [2], a brief summary 

of the thermal behavior of steel members and characterization of the furnace environment 

are provided here. In this report, the experimental set up and the fire resistance tests are 

described first. Then experimental results collected at several locations in the trusses and 

furnace are presented. A brief statistical analysis of member temperatures measured 

during the full-scale tests is then provided. Finally, numerical results from the heat 

transfer model are compared against experimental data. A quantitative estimate of the 

agreement between the experimental and simulation results is then provided. Future work 

will focus on estimating uncertainties in member temperatures.

2 Description of the Fire Resistance Tests

As part of its investigation into the collapse of the World Trade Center (WTC) towers, 

NIST conducted four fire resistance tests of composite floor systems (two full-scale and 

two reduced-scale). These tests were carried out by Underwriters Laboratories (UL) by 

subjecting composite floor systems to the heating profile (time-temperature curve) 

specified in the ASTM E-119 standard in a test furnace.

The tests were conducted in both restrained and unrestrained condition to study the 

performance of the floor systems under the standard fire exposure. Note that in a 

restrained test, thermal expansion at the supports of a load carrying element due to 
7

exposure to elevated temperatures in fire is resisted by forces external to the element. On 

the other hand, an element is free to expand and rotate at its supports in an unrestrained 

test. In this section, a brief description is provided about the test specimens, test furnace, 

instrumentation, and test procedure pertaining to the two full-scale tests. 

2.1 Test Specimen

The floor system for the WTC towers consisted of a lightweight concrete floor slab acting 

compositely with steel trusses bridging between the building’s core columns and exterior 

wall columns. The main composite trusses, which were used in pairs, spanned either 60 

ft.1 (18.3 m) or 35 ft. (10.7 m) Steel double angles formed the top and bottom chords of 

the trusses while round bars were used for the web diagonals. Additionally, the floor 

system included bridging trusses perpendicular to the main trusses. 

Passive fire protection was provided by spray applied fire resistive material (SFRM), 

commonly referred to as “fireproofing,” applied directly to the steel members of the 

trusses. Fireproofing thickness was specified to be 0.75 in. (19 mm). Figure 1(a) shows a 

cut-away of the floor system illustrating the main trusses, bridging trusses, metal deck, 

and concrete floor slab used in the WTC towers. Figure 1(b) shows a cross-sectional view 

of the basic configuration of the truss floor system.

Trusses used in the WTC towers were fabricated from ASTM A242 grade steel [2]. Since 

this steel was no longer available, ASTM A572 Grade 50 steel was found to be an 

acceptable substitute by NIST based on a comparison of thermal and mechanical 

properties of both steels. For the full-scale fire resistance tests, ASTM A572 Grade 50 

steel test specimens were constructed to model the 35 ft. (10.7 m) span floor trusses. The 

deck consisted of 11⁄2 in. (38.1 mm). deep 22 gauge galvanized sheet metal floor units. 

Compressive strength of the concrete floor slab was specified to be 3000 psi (20.7 MPa). 

The sprayed fire resistive material used on the test assemblies was BLAZE-SHIELD

Type DC/F, which is manufactured by Isolatek International. The SFRM was applied to a 

thickness of 3⁄4 in. (19 mm). with close tolerances of approximately ±1/16 in. (1.6 mm).

The test specimen comprised two main trusses, two bridging trusses, and a floor deck

(Figure 2).

Two full-scale standard fire resistance tests were conducted at the Underwriters 

laboratories fire testing facilities:

1. Test 1: restrained test condition.

2. Test 2: unrestrained test condition.

 1 The policy of NIST is to use the International System of Units (metric units) in all publications. In this 

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

the discipline.
8

2.2 Furnace

The Underwriters Laboratories of Canada (ULC) fire testing facility in Toronto, Canada 

was used for the fire resistance tests. ULC has a furnace with nominal dimensions of 35 

ft. (10.7 m) by 14 ft. (4.3 m). The inside furnace depth is approximately 5 ft. (1.5 m) Test 

assemblies were constructed so that the top of the floor assembly was nearly flush with 

the top of the reaction frame. The combustion system consisted of 80 individual floor 

mounted burners, arranged in an evenly spaced grid of 8 rows with 10 burners in each 

row for producing furnace temperatures that are as uniform as possible.

Figure 1(a). Floor system of the WTC towers [3].

Figure 1(b). Schematic of floor system viewed along the main steel trusses (not to scale)

[3].
9

2.3 Instrumentation

The furnace temperature was measured by means of 24, 16-gauge, Type K 

thermocouples, sheathed in Inconel pipe symmetrically located in the furnace chamber, 

as required by the ASTM E-119 standard. The thermocouples were positioned about a 

foot below the bottom surface of the exposed floor slab. In addition to the furnace 

thermocouples, Wickstrom plate thermocouples and aspirated thermocouples were also 

used to monitor furnace temperatures. In addition, heat flux gauges were also used to 

characterize the thermal environment inside the furnace.

A total of 88 thermocouples were used to measure steel temperatures at various locations 

along the trusses [2]. Figure 2 shows thermocouple locations in the main trusses.

Instrumentation was also added to measure temperature on the unexposed surface of the 

concrete slab at several locations [2].

Figure 2: Thermocouple locations in the main truss in the UL tests [1].

Position at a cross section

Locations 

chosen

North Main 

Truss

South Main 

Truss

Bridging Trusses
10

2.4 Test Procedure

The tests were conducted according to the procedure specified in the ASTM E119

standard [4]. The ASTM E119 standard prescribes a standard fire exposure for comparing 

the test results of building construction assemblies. For the tests of floors and roofs, the 

full design load is applied to the specimen and the standard fire exposure is applied to the 

underside of the specimen. The assembly is evaluated for its ability to contain the fire by 

limiting flame spread and by limiting heating of the unexposed surface while maintaining 

the applied load. Acceptance criteria are provided in terms of time to reach the stipulated 

average and maximum temperatures both on the unexposed surface and in steel members. 

The assembly is given a rating, expressed in hours, based on these acceptance, or 

endpoint, criteria.

The average of 24 thermocouple readings of furnace temperatures for both tests is shown 

in Figure 3. Also shown is the prescribed ASTM E119 curve. Although there are 

fluctuations in temperatures recorded by the furnace thermocouples especially during the 

initial periods of the test, the mean temperatures seem to follow the ASTM E119 curve 

very well. Note that the furnace temperatures for Tests 3 and 4 are also shown but are not 

discussed in this report.

Figure 3: Average furnace temperatures and the specified ASTM E119 curve [1].

The details of the test assembly preparation as well as the loading procedure are 

described in detail in Ref. [2]. 

Fire Resistance Testing of WTC Floor System 

Average Furnace Control Temperature Comparison

0

200

400

600

800

1000

1200

1400

1600

1800

2000

50 100 150 200

Time (minutes)

Temperature (deg. F)

ASTM E119

ULC Restrained Assembly (Test 1)

ULC Unrestrained Assembly (Test 2)

UL Restrained Assembly (Test 3)

UL Restrained Assembly (Test 4)
11

3 Experimental Results

3.1 General Observations of Tests

In Test 1, loud noises (described as “reports” in [2]) were heard beginning roughly fifteen 

minutes after the start of the test (see Appendix C for a log of observations throughout 

these tests). These loud reports continued and were often accompanied by noticeable 

bulging of the metal deck and the dislodging of deck fireproofing. It is believed that these 

loud noises were associated with spalling of the concrete on the underside of the slab. 

Reports were also heard during Test 2, of the 35 ft. (10.7 m) unrestrained specimen, but 

were not as loud as those observed during Test 1. Post-test inspection of the top of the 

concrete slab showed very little cracking. Since there was a gap between the concrete 

slab and the test frame, thermal expansion of the slab did not produce a state of stress that 

resulted in significant and sudden cracking and crushing similar to this test. Appendix C

and D provides a log of the reports observed during the tests.

3.2 Furnace Thermal Environment

In addition to the 24 sheathed thermocouples that were used to measure and monitor 

furnace temperatures, furnace temperatures were also measured with plate and aspirated 

thermocouples. Plate thermocouples are used to measure gas temperatures with a surface 

similar in terms of emissivity, size, and orientation to the specimen surface. Aspirated 

thermocouples help decrease the magnitude of error by reducing the radiative exchange 

between the thermocouple and its surroundings. In fact, temperatures measured by 

aspirated thermocouples are closer to the true gas temperatures [5]. 

Figures 4 and 5 show the temperatures measured by plate and aspirated thermocouples 

respectively. It is evident from these figures that the thermal environment in these two

tests was similar.

Since thermal radiation is the dominant form of heat transfer, especially at elevated 

temperatures, heat flux measurements provide a better indication of how quickly a 

specimen will heat than gas temperatures. For Test 2, radiometer readings were not 

available, due to damage to the radiometers. For Test 1, the radiative heat flux was about

60 kW/m2 at the end of 104 minutes, after which the heat flux data showed wide 

fluctuation and data became unreliable (Figure 6). A dip in the curve at about 50 minutes 

into the test indicates that the probes recorded a lower radiative heat flux. A change in 

“view factors” resulting possibly from concrete spalling around that time could have 

contributed to a lower value of heat flux recorded (see the log of the test in Appendix C).
12

0

200

400

600

800

1000

1200

0 50 100 150

Temperature, degree C

Time, minutes

Test 1 Plate TC

Test 2 Plate TC

Figure 4: Temperatures measured by plate thermocouples in tests 1 and 2 [2].

0

200

400

600

800

1000

1200

0 50 100 150

Temperatures, degree C

Time, minutes

Test 1 Aspirated TC

Test 2 Aspirated TC

Figure 5: Temperatures measured by aspirated thermocouples in tests 1 and 2 [2].
13

0

40

80

120

160

200

240

0 50 100 150

Radiative Flux (kW/m2)

Time, minutes

Radiative Heat Flux

Figure 6: Radiative heat flux in Test 1 [2].

3.3 Unexposed Concrete Slab Surface Temperatures

The average temperature of the top (unexposed) surface of the floor assemblies is plotted 

in Figure 7. It is observed that the unexposed surface temperatures of both test assemblies 

were similar prior to the onset of significant concrete crushing and spalling at around 50 

minutes in Test 1. Thereafter, the average unexposed surface temperature in Test 1 

steadily increased in comparison to that in Test 2 as shown in Figure 7. Two factors could 

have contributed to higher unexposed surface temperatures in Test 1: a) Spalling of 

concrete slab at several locations thereby offering less resistance to heat flow through the 

slab and b) effect of varying moisture level on concrete thermal conductivities.

Figure 7: Average concrete slab temperature on unexposed surface in Tests 1 and 2.
14

3.4 Steel Temperatures

Figures 8(a) through 8(d) show a comparison of average temperatures at the bottom 

chord for both tests at four locations e.g., A, C, I, and E (refer to Figure 2 for 

thermocouple locations). For the bottom chords, there were four thermocouples (two on 

each angle at the bottom chord). Figure 8 shows average temperatures at each angle in the 

bottom chord (as measured by TC 3 & 4 and TC 7 & 8 respectively). These plots show 

that the heating trends in the bottom chords are similar in both tests, until about 60 

minutes into the test when data from Test 1 showed an abrupt increase in recorded 

temperatures, possibly due to dislodgement of SFRM as a result of concrete spalling 

(Figure 8(d)). However, the plots also show a wide variability in temperatures recorded 

by thermocouples on two angles at the bottom chord (e.g., Fig. 8(a): “Test 1, Section A, 

TC 3 & 4”, and “Test 1, Section A, TC 7 & 8”).

Figure 8 (a) through (d): Comparison of two average temperatures in the bottom chord in 

Test 1 and 2 [2].

(a) (b)

(c) (d)
15

Figures 9 (a) through (d) show a comparison of average temperatures recorded by the 

thermocouples at the top chord of both the test specimen. Individual thermocouple 

temperatures have been shown when data for both the thermocouples at the top chord 

were not available for a particular location along the main truss. It is possible that 

dislodgement of SFRM took place at around 60 minutes in Test 1 near locations I, C, and 

E. For Test 2, it appears that such possible dislodgement took place in the vicinity of 

location C at about 70 minutes into the test. Heating trends for both the tests were 

generally similar. Location E shows the most discrepancy in measured temperatures 

between Test 1 and 2 after about 60 minutes into the test.

Most thermocouples at the top chord recorded maximum temperatures of less than 

400  ̊C after 2 hours of heating in both the tests. Some thermocouples recorded abrupt 

temperature increases, possibly due to the dislodgement of SFRM. A quantitative 

comparison between thermal results in the two tests is provided later in this section.

Figures 10(a) through (d) show a comparison of average temperatures in the two tests for 

four locations at mid-web along a main truss. Although time-dependent temperatures at 

location A for both the tests were close to each other (until about 100 minutes into the 

test for Test 1 when thermocouple data became unreliable and showed wide fluctuations 

in temperatures), the average temperatures in Test 1 were higher than those in Test 2 for 

locations C, I, and E.

Figure 9 (a) through (d): Comparison of two average temperatures in the top chord in Test 1 and 

2.

(a) (b)

(c) (d)
16

Figures 10 (a) through (d): Comparison of two average temperatures in the mid-web

regions for Test 1 and 2.

3.5 Variability in Experimentally Recorded Temperatures

There is significant variation in recorded temperatures along the top chord, (a) in the 

same position (A 1 and A 5 in Test 1) (Figure 11(a)), (b) in the same location at two main 

trusses (A 1 and B 1) (Figure 11(b)), and (c) at different locations along a truss length 

(A1 and E1) (Figure 11(c)). Such wide difference in recorded temperatures indicates the 

need to determine the variability in the time-dependent, non-uniform temperature 

distribution in steel members. It is important to determine the factors that could possibly 

contribute to this variability. Some of these are: a) non-uniform heating as a result of 

varying gas temperatures based on locations, b) time-dependent change in “view factors” 

due to concrete spalling or SFRM buildup, c) change in net convective heat transfer due 

to changes in local gas temperatures near the vicinity of the thermocouples. The second 

phase of this research will attempt to compute uncertainties in temperatures and compare 

with those obtained in measurements.

(a) (b)

(c) (d)
17

0

200

400

600

800

0 20 40 60 80 100 120 140

Temperature, 0C

Time, minutes

Location A

A 5

A 1

0

200

400

600

800

0 20 40 60 80 100 120 140

Temperature, 0C

Time, minutes

Locations A and E

A 1

E 1

Figure 11: Comparison of experimentally recorded temperatures in Test 1 for top chord 

locations.

0

200

400

600

800

0 20 40 60 80 100 120 140

Temperature, 0C

Time, minutes

Locations A and B

A 1

B 1

(a) (b)

(c)
18

Table 1 lists the means and standard deviations of the average measured temperatures at 

lower chord, mid-web, and upper chord locations for the two main trusses for two 

different times (60 minutes and 120 minutes into the test). Recall that the data were 

recorded at 10 locations along the two main trusses (A, G, C, I, and E in the north main 

truss and B, H, D, J, and F in the south main truss; see Figure 2). At each location, two 

thermocouple readings were available at the top chord and mid-web, while four readings 

were available at the bottom chord. Average of these temperature readings at the top 

chord, mid-web, and the bottom chord were calculated at each of the 10 locations along 

the two main trusses. Then, the overall mean temperatures and the standard deviations at 

the top chord, mid-web, and the bottom chord were computed considering these 10 

average temperatures for both tests. In computing these data, anomalous thermocouple 

readings were excluded. 

These data are plotted in Figure 12. The mean temperature data show that the web 

members heated up the fastest for both tests. This is expected since a) heating of the web 

can be considered to be uniform and the influence of the floor slab is presumed to be 

minimal b) the exposed surface area/volume ratio of web members are greater than that 

of the chord members. The pattern of heating for the two tests shows similar trends. Also, 

the mean temperatures in the lower chord, mid-web, and upper chord of the truss in both 

the tests are reasonably close. Clearly, Test 1 showed wide variability in mean 

temperatures (see standard deviations for Test 1 in Table 1). This is likely due to the 

concrete spalling as discussed earlier. This event changed the subsequent thermal profile 

in the truss, contributing to variability in recorded temperatures. The upper chord mean 

temperature at 120 minutes was about 33  ̊C higher in Test 1 than in Test 2.

Table 1. Means and standard deviations of average temperatures at three locations on the 

main truss in tests 1 and 2 at 60 minutes and 120 minutes into the fire.

Location on

main truss μ, 

0

C σ,

 0C μ, 

0

C σ,

 0C μ, 

0

C σ,

 0C μ, 

0

C σ,

 0C

Lower Chord 445.7 81.8 446.4 54.3 773.3 69.8 772.0 37.9

Mid Web 600.4 35.0 576.5 36.6 860.3 70.3 872.2 42.9

Upper Chord 161.3 35.4 162.9 29.0 358.1 108.6 325.4 39.5

60 minutes 120 minutes

Test 1 Test 2 Test 1 Test 2
19

Figure 12: Mean temperatures and two standard deviations at 3 locations on main trusses 

in tests 1 and 2 at 60 minutes (a, b) and 120 minutes (c, d) into the tests.

The measured temperatures at 4 locations show that the difference between the average 

temperatures at the top chord and mid-web positions was: 218  ̊C at 30 minutes and 

502  ̊C at 2 hours in Test 1 respectively and 200  ̊C at 30 minutes and 546  ̊C at 2 hours

in Test 2 respectively. The difference between the average temperatures at the top chord 

and bottom chord positions were: 95  ̊C at 30 minutes and 415  ̊C at 2 hours in Test 1 

respectively and 105  ̊C at 30 minutes and 446  ̊C at 2 hours in Test 2 respectively. 

Additionally, the measured temperatures at 4 locations also show that the maximum 

differences between the average temperatures along the member length were: 

(a)

(a) (b)

(c) (d)
20

• 34  ̊C and 76  ̊C at the mid-web at 30 minutes and 90 minutes respectively in Test 

1 and 85  ̊C and 79  ̊C at the mid-web at 30 minutes and 90 minutes respectively 

in Test 2.

• 74  ̊C and 124  ̊C at the bottom chord at 30 minutes and 90 minutes respectively 

in Test 1 and 38  ̊C and 167  ̊C at the bottom chord at 30 minutes and 90 minutes 

respectively in Test 2.

• 46  ̊C and 128  ̊C at the top chord at 30 minutes and 90 minutes respectively in 

Test 1 and 38  ̊C and 77  ̊C at the top chord at 30 minutes and 90 minutes 

respectively in Test 2.

These data indicate that the time-dependent variation of temperatures along the length of 

a member must be considered in conducting structural analysis during fire tests. One 

conclusion that can be drawn from the above discussion is that concrete spalling 

contributes significantly to the overall uncertainty in steel member temperatures. 

Variabilities in member thermophysical properties, fire/member heat exchange 

parameters, and gas temperatures must be considered to determine the overall 

uncertainties in member temperatures.

4 Heat Transfer Modeling

A finite element model of the composite floor system described in section 2.1 was 

developed using the ANSYS finite element software. The dimensions of the model were 

taken from Ref. [2]. The purpose of the heat transfer model is to simulate the transient 

heat diffusion in the fire-exposed composite floor system. The heat transfer model was

used to compare model predictions of temperatures at a few locations in the floor system 

with experimental measurements. Thermal data predicted by this validated heat transfer 

model will be later used in a structural analysis model for predicting the time-dependent 

behavior of the floor system due to both thermal and gravity loads.

In this section, the following topics are covered:

1) Details of the methodology used to develop the heat transfer model

2) Description of input data used for thermophysical properties of materials in the 

floor system and the boundary and initial conditions

3) A brief description of the approach used to interpolate temperature values (which 

are compared with experimental measurements)

4.1 Methodology

A finite element model of the composite floor system shown in Figure 2 was constructed 

by including the following:
21

a. One steel main truss (without end supports) including the top chord, web 

diagonals, and the bottom chord.

b. Concrete slab (above the main truss) spanning the entire truss length and half the 

furnace width of 2.15 m (see Figure B1 in Appendix B).

Note that only one main truss shown in Figure 2 was modeled in this study. Bridging 

trusses were not included. The spray applied fire resistive material (SFRM) was 

considered to be uniformly applied over the steel truss (no fireproofing was applied on 

the top surface of the top chord). Additionally, the exposed surface of the slab was 

assumed to have an overspray of SFRM of about 19 mm. Note that the profiled steel 

metal deck was not specifically modeled since the metal deck is only 0.9 mm thick, while

the slab is about 130 mm thick. Moreover, Lamont’s [6] results did not show a significant 

difference in the temperature prediction in the slab when a metal deck was specifically 

modeled in comparison with the case when the metal deck was not included.

The finite element model of the composite floor assembly was developed using 3-D solid 

thermal elements (SOLID 70 element in ANSYS), which have eight nodes with a single 

degree of freedom (temperature) at each node. The truss upper chords and lower chords 

were connected with the web members using conduction link elements (Link 33 element 

in ANSYS). These are uniaxial elements that can be used to model the transient heat 

transfer between the end nodes. Use of link elements allows the diagonals to be meshed 

independently of the upper and lower chords and resulted in the minimum possible 

volumes in the model. Figures B1 and B2 (Appendix B) show the FE (finite element)

model in detail. Note that adequate numbers of elements had to be used through the depth 

of each material to ensure that at least two free nodes existed at each section and that 

steep thermal gradients, if any, are handled correctly in the model.

During fire exposure, heat exchange occurs by both convection and radiation. At higher 

temperatures, radiation plays a more dominant role in thermal exchange. The convective 

heat exchange is addressed by the Newtonian heating/cooling mechanism. The radiative 

heat exchange is defined by the Stefan-Boltzmann law, which includes a parameter called 

“view factor”. The view factor is defined by the proportion of the radiation that leaves a 

surface (e.g., surface A) that strikes another surface (e.g., surface B). It can assume 

values from 0 to 1.

In order to simulate the furnace heating in the UL tests, the exposed surfaces of the 

composite floor system were subjected to convective and radiative heat exchanges. In 

order to accomplish this, surface effect elements were generated in ANSYS and overlaid 

on the exposed surfaces of the solid elements. This is because only one type of thermal 

surface load can be applied to a given entity in ANSYS. For example, one cannot apply 

both a flux and a convection boundary condition directly to the faces of an element. 

Surface effect elements allow the use of multiple surface loads to a given element face. 

Two types of surface effect elements were created: one for the convective heat exchange 

and the other for the radiative heat exchange. All of these surface effect elements were 

created with a particular reference node. For the convective surface elements, the fluid 
22

bulk temperature was designated as that of the reference node and appropriate convective 

heat transfer coefficients were used (see the next section). For the radiative heat 

exchange, surface effect elements were created with a reference node. An imposed 

temperature constraint (e.g., furnace heating following the ASTM E119 curve) was 

specified on this node. A resultant emissivity value was specified while creating these 

surface effect elements. The resultant emissivity includes proper emissivity values of the 

receiving and emitting surfaces. View factor effects in the radiation heat transfer were not 

considered, since the furnace was not explicitly modeled in this study and therefore, view 

factor values were set to 1 for all elements. 

4.2 Thermophysical Properties, Boundary and Initial Conditions

Thermophysical properties for materials used in the model (e.g., steel, concrete, and 

SFRM) are taken from Ref. [7] and are listed in both tabular and graphical forms in 

Appendix A. Note that temperature-dependent values of materials properties were used. 

The three thermophysical properties that are needed for thermal analysis are: density, 

specific heat, and thermal conductivity.

Proper boundary conditions are needed to model the heat transfer to/from the floor 

system both at the exposed and unexposed surfaces (e.g., at the bottom and top surfaces

of the slab). For the convective surface effects elements at the exposed surface, a 

convective heat transfer coefficient is specified. A value of 25 W/m2

/K has often been 

used [8]. Lamont et al. [6] used a base value of 23 W/m2

/K, with minimum and 

maximum values of 5 W/m2

/K and 30 W/m2/K in their sensitivity analysis. Lamont et al.

suggested using a lower value of this heat transfer coefficient when the profiled metal 

deck is not explicitly modeled. This is because the profiled metal deck is known to 

separate from the rest of the concrete slab during progression of a fire event, thereby 

creating a gap that reduces the convective heat transport at the exposed surface of the 

slab. In this analysis, a value of 15 W/m2/K was used for the convective heat transfer 

coefficient since the profiled metal deck was not modeled this study.

The effective emissivities for radiative heat exchange at the exposed surface depend on 

many factors such as flame emissivity, compartment walls, etc. For the surface emissivity 

of the steel member, a value of 0.7 is used. At the exposed surface of the slab, a value of 

0.6 is used for effective emissivity as recommended by Lamont [6]. This value is slightly 

different from that of steel and was used as the metal deck was not explicitly modeled in 

this study. This is because of the air gap that forms at the interface between metal deck 

and concrete slab. Lamont found very good agreement between modeled and measured 

concrete slab temperatures when this value of emissivity was used. 

Note that adiabatic boundary conditions were maintained at each end of the model in the 

longitudinal direction of the truss. Average time-temperature values for different 

segments on the unexposed surface of the concrete slab as determined from recorded 

thermocouple data during the two fire tests were used as Dirichlet boundary conditions 

(time varying) for nodes on the unexposed surface of the concrete slab. Note that this 
23

boundary condition is the only difference between Test 1 and Test 2 from the thermal 

modeling perspective. All nodes in the model were assigned an initial temperature of 

25  ̊C. The input temperatures to the model are the ASTM E119 time-temperature curve

assigned to the reference node as described in the previous section. This essentially

assigns the furnace heating to this reference node. All exposed surfaces exchange 

convective and radiative heating with this node.

4.3 Interpolation of Temperatures

The numerical model predictions can be compared with measured time-temperature data 

at locations A, C, E, and I along the main truss (see Figure 2). These comparisons are 

performed at the upper chord, mid-web, and the lower chord for these four locations.

However, often there is not a node at the exact location where a thermocouple was 

positioned. In order to obtain a simulated temperature at the measurement point, 

interpolation of temperatures was done as described in the following paragraphs. 

Figure 13. Schematic illustration of the averaging approach used to interpolate

temperatures.

The model predictions of temperature at a given location along the truss (e.g., A, C, E, or 

I in Fig. 2) are obtained as shown schematically in Figure 13. For example, temperature 

at the position of interest (e.g., P at the bottom chord in Figure 13) is obtained by taking a 

weighted average of temperatures of all the FEM nodes available nearest to the point P 

on either side. In other words, first average temperatures at X and Y are obtained by 

finding the mean temperatures of all the nodes at X and Y (e.g., through the entire 

sections at X and Y). Then the temperature at the point of interest (P) is estimated by 

finding a weighted average of computed mean temperatures at X and Y using the proper 

distance of P from both locations X and Y. A script was written in ANSYS parametric 

design language (APDL) to output these temperatures as functions of time for the desired 

locations along the truss for upper chord, mid-web, and lower chord positions. These

Y

X

P
24

temperatures were then compared with measured temperatures available for these 

locations.

5 Evaluation of the Heat Transfer Model

In this section, average computed time-temperature data at four selected locations along 

(A, C, E, and I) the North Main Truss are compared against the experimentally obtained 

results in both the tests. As mentioned earlier, these measurements were taken at bottom 

chord, mid-web, and upper chord positions at five locations along each main truss. 

Whenever experimental data showed anomalous temperatures, readings from the South 

Main Truss for the identical location have been included in plots for comparison. 

Additionally, erroneous thermocouple data have not been included. Note that erroneous 

thermocouples are listed in the log of experimental data as reported in ref [2].

5.1 Test 1

Figures 14, 15, and 16 show comparisons between ANSYS predicted mean temperatures 

and measured temperatures at these four locations along the north main truss for positions 

at bottom chords, mid-webs, and top chords (see Figure 2 for labeling of positions).

Figure 14: Time-temperature plots of the thermocouple data recorded for the bottom 

chord at four locations on the north main truss in UL Test 1. Also shown are the average 

computed temperatures obtained from ANSYS.
25

Figure 15: Time-temperature plots of the thermocouple data recorded for the mid-web

position at four locations on the north main truss in UL Test 1. Also shown are the 

average computed temperatures obtained from ANSYS.

Table 2 shows a listing of measured and ANSYS computed temperatures for four 

different times during the test at the lower chord, mid-web, and upper chord for the four 

locations along the truss. Additionally, percentage differences of computed temperatures 

(in absolute scale e.g., K) from experimental values are also listed. It is evident that the 

agreement is good for both bottom chord and mid-web locations except for the location I

at the bottom chord, where the difference between the measured and computed 

temperatures was as high as 12%. Such disagreement is also present in Test 2 as will be 

discussed later. 
26

Figure 16: Time-temperature plots of the thermocouple data recorded for the top chord 

position at four locations on the north main truss in UL Test 1. Also shown are the 

average computed temperatures obtained from ANSYS.

A loud report was heard at 45 minutes (see Appendix C) when concrete spalling 

occurred, resulting in dislodgement of fireproofing from the top chord (see Ref [2]). This 

resulted in an abrupt rise in steel temperatures as shown by the experimental plots for the 

top chord (Figure 16). These plots show that the thermal model has been set up correctly 

since the initial rate of heating matches closely with the rate shown in the measured 

curves. Moreover, Table 2 shows that the agreement is within 10% for most mid-web and

bottom chord locations. Another observation can be made from Table 2: the 

discrepancies between measured and computed temperatures for mid-web and lower 

chord positions are somewhat large during the initial periods of the test. This discrepancy 

may be attributed to the fact that the model includes only one truss and transient effects 

due to non-uniform heating and heat exchanges between members such as bridging 

trusses, and the south main truss were not included here. This discrepancy diminishes at 

later times into the test as gas temperatures became more uniform in the furnace. 
27

Table 2: A comparison of experimental and predicted temperatures at four times for

Test 1.

Test1

Mid Web Temperatures in C (% diff in absolute scale (K))

Time, min Exp (A) ANSYS (A) % diff Exp (C) ANSYS (C) % diff Exp (I) ANSYS (I) % diff Exp (E) ANSYS (E) % diff

30 331 285 -7.7 299 289 -1.7 320 285 -5.8 332 280 -8.6

60 647 626 -2.2 656 633 -2.4 618 625 0.7 621 617 -0.4

90 792 820 3.5 846 826 -1.8 770 817 4.5 775 811 3.4

120 - 929 - - 934 - 899 927 2.4 933 923 -0.9

Bot Chord

Time, min Exp (A) ANSYS (A) % diff Exp (C) ANSYS (C) % diff Exp (I) ANSYS (I) % diff Exp (E) ANSYS (E) % diff

30 245 237 -1.7 180 200 4.5 172 198 5.8 195 229 7.4

60 565 578 1.5 477 502 3.4 417 498 11.7 461 561 13.6

90 752 782 2.9 747 709 -3.7 628 704 8.5 735 765 3.0

120 929 903 -2.1 - 841 - 781 838 5.3 768 889 11.6

Top Chord

Time, min Exp (A) ANSYS (A) % diff Exp (C) ANSYS (C) % diff Exp (I) ANSYS (I) % diff Exp (E) ANSYS (E) % diff

30 92 97 1.3 100 101 0.5 86 102 4.3 132 100 -7.9

60 209 196 -2.7 - 203 - 169 203 7.5 - 151 -

90 382 278 - - 284 - 254 284 5.8 - 282 -

120 574 348 - - 353 - 336 353 2.8 - 351 -

5.2 Test 2

Table 5-5 of Ref. [2] indicates that it took 76 minutes of heating for average steel 

temperatures in Test 2 to reach 593.3  ̊C (time to reach End-point Criteria in ASTM E 

119 test), while this duration was 66 minutes for Test 1. These data indicate that the 

heating rate in steel members was slower in Test 2 that it was in Test 1. This is also 

evident in Figures 8, 9, and 10, where average temperatures recorded in the two tests 

along a main truss are compared. Factors such as heat loss could play a role in somewhat 

reduced steel member heating rates seen in Test 2. Moisture level in concrete can also 

play a role as a concrete slab with higher moisture content is expected to be a more 

effective heat sink.

Figures 17 through 19 compare measured and predicted steel temperatures at lower 

chord, mid-web, and upper chord positions along the length of the truss. It can be noticed 

that measured bottom chord temperatures for most locations are slightly lower than 

predicted temperatures. The agreement at the bottom chord is within 10%. The agreement 

at mid-web position is generally within 5%. For upper chord, the agreement is within 4% 

except at location I. This discrepancy can also be seen in Table 3, where measured and 

predicted temperatures are listed at four different times during the fire test.
28

As seen in Test 1, the measured temperatures at location I in Test 2 are substantially 

lower than those predicted by the model. The reason behind this is not quite clear. One 

possible explanation is that since the bridging truss intersects the main truss at location I

(see Figure 2), it could act as heat sink. It may be noted that although the bridging truss 

was heated during the tests, it could still act as a heat sink because of its thermal mass.

Since the bridging truss was not included in the model, this discrepancy could not be 

addressed. In the future, the computational model will include the entire floor system 

(e.g., both main trusses and bridging trusses).

Figure 17: Time-temperature plots of the thermocouple data recorded for the bottom 

chord at four locations on the north main truss in UL Test 2. Also shown are the average 

computed temperatures obtained from ANSYS.
29

Figure 18: Time-temperature plots of the thermocouple data recorded for the mid-web at 

four locations on the north main truss in UL Test 2. Also shown are the average 

computed temperatures obtained from ANSYS.
30

Figure 19: Time-temperature plots of the thermocouple data recorded for the top chord at 

four locations on the north main truss in UL Test 2. Also shown are the average 

computed temperatures obtained from ANSYS.
31

Table 3: A comparison of experimental and predicted temperatures at four time levels for 

Test 2.

Test2

Mid Web Temperatures in 0

C ((% diff in absolute scale (K))

Time, min Exp (A) ANSYS (A) % diff Exp (C) ANSYS (C) % diff Exp (I) ANSYS (I) % diff Exp (E) ANSYS (E) % diff

30 337 284 -8.6 308 289 -3.2 276 259 -3.1 252 255 0.6

60 643 626 -1.9 621 633 1.4 569 566 -0.4 543 559 1.9

90 798 819 2.0 793 825 3.1 738 753 1.5 720 746 2.7

120 943 929 -1.1 879 934 4.7 884 867 -1.4 852 862 0.9

Bot Chord

Time, min Exp (A) ANSYS (A) % diff Exp (C) ANSYS (C) % diff Exp (I) ANSYS (I) % diff Exp (E) ANSYS (E) % diff

30 210 211 0.3 211 185 -5.3 173 181 1.7 195 204 1.9

60 495 530 4.5 498 469 -3.7 399 459 9.0 447 512 9.0

90 705 739 3.5 769 676 -8.9 602 664 7.1 646 720 8.0

120 808 868 5.6 835 810 -2.3 768 798 2.9 791 850 5.6

Top Chord

Time, min Exp (A) ANSYS (A) % diff Exp (C) ANSYS (C) % diff Exp (I) ANSYS (I) % diff Exp (E) ANSYS (E) % diff

30 104 97 -2.0 98 100 0.6 66 101 10.4 103 99 -1.0

60 189 194 1.0 186 200 3.2 116 200 21.8 196 198 0.4

90 261 275 2.7 - 280 - 203 280 16.2 280 277 -0.4

120 346 345 -0.2 - 346 - 285 346 11.0 369 344 -3.9
32

5.3 Effect of SFRM Dislodgement

As shown in Figure 9(a) through 9(d), the top chord temperatures for the two tests show 

similar heating trends on average. However, the reduction in the thickness of SFRM as a 

result of dislodgement possibly resulted in the sudden increase in the recorded 

temperatures for a few locations for Test 1. The aim in this section is to demonstrate an 

approach for estimating the effect of SFRM dislodgement at specific locations.

The analysis of Test 1 showed significant effects of concrete spalling on temperatures at 

the top chord (Figure 9). ANSYS simulation was interrupted according to the timeline 

corresponding to the loud reports indicated in the log of observations [Appendix A and 

B] and 2 ft. (0.61 m) to 3 ft. (0.91 m) of SFRM in length were removed from the model at 

these locations. ANSYS simulation was then restarted and continued until the occurrence 

of the next loud report. A trial and error approach was followed to determine the range of 

possible amounts of SFRM dislodgement presumed to occur at each location. A 

schematic of SFRM dislodgement from the model is shown in Figure 20.

Figure 20: A schematic of the SFRM removal at locations A, C, I, and E in Test 1.

For Test 1, at 3800 s, 0.5 in. (12.7 mm) of SFRM was removed at location A and 0.75 in. 

(19 mm) of SFRM was removed at locations I and E. Simulation was then restarted and 

continued until 4500 s, when 0.75 in. (19 mm). of SFRM was removed at location C and 

then simulation was restarted. Figure 21 shows a comparison of computed and 

experimental data at the four locations along the top chord in Test 1. For locations C and 

E, computed results are also shown for the hypothetical case of SFRM remaining intact. 

As expected, the temperature plots (for this case of SFRM remaining intact) continue 

along the “ANSYS simulated, SFRM removed” curves before abrupt heating took place 

during the onset of dislodgement of SFRM (see dashed plots for ANSYS simulation for 

locations C and E in Figure 21).

SFRM Removed

C

A

I E
33

Figure 21: Comparison of numerical results and experimental data for upper chord 

temperatures in Test 1 with SFRM removal.

It can be seen that the average computed temperatures match reasonably well with the 

thermocouple data at location A. At location C, the computed rate of heating was higher 

than that of the experimental plot. This is true until about 75 minutes when there is an 

abrupt temperature increase possibly due to dislodgement of SFRM from steel. There is a 

reasonable match between the computed and experimental data at location E except 

toward the end, when the experimental data show a dip in temperature followed by a 

gradual increase toward the end. This analysis shows that it is possible to estimate the 

extent of SFRM dislodgement from steel members during a fire exposure by conducting a 

controlled and iterative modeling exercise such as the one shown here.

(b)

(c)
34

6 Summary

An analysis of the measured temperatures of the main trusses in the two full-scale fire 

resistance tests showed that the steel web members heated faster than the bottom and top 

chord members and at the end of the test, the steel web members recorded the highest 

temperatures. On the whole, the restrained test (Test 1) showed higher variability in 

recorded steel temperatures than that in the unrestrained test (Test 2). This is because 

there was less cracking and spalling of concrete in the unrestrained test. Additionally, top 

chord temperatures in the restrained test showed wide variability especially toward the 

end of the fire tests. This conclusion can be drawn based on large values of standard 

deviations obtained.

An ANSYS heat transfer model was developed with appropriate thermophysical 

properties and model parameters (e.g., boundary conditions, initial conditions etc.) to 

predict the thermal behavior of the structural members. The heat transfer model was 

evaluated by comparing the time-dependent evolution of measured mean temperatures at 

lower chord, mid-web, and upper chord positions along four specific locations of the 

main trusses with average computed temperatures. For the bottom chord and mid-web

locations, the agreement between measured and ANSYS temperatures was within 15%

for both tests. For the top chord, the agreement was within 10% until about 60 minutes 

for Test 1, when concrete spalling occurred. For Test 2, the agreement at the top chord 

was within 10% for all locations except at location I, where the disagreement was 

substantial. The cause of this disagreement is not clear but it could possibly be attributed 

to the fact that the bridging truss acted as a thermal sink and that it was not included in 

the model. 

Measured steel temperatures during fire tests indicated that the variation in temperatures 

among top chord, web diagonal, and bottom chord members was substantial. Also, there 

was large temperature variation along the length of the steel truss, which varied as tests 

progressed. These observations suggest that both initial conditions (variations in 

composition and geometry) and time varying boundary conditions (including the 

handling of material loss/transformation) are necessary to consider in order to predict the 

temporal and spatial variations of steel temperatures.

The present study provided the following information:

• A brief statistical evaluation of the thermal data recorded during the two full scale 

fire tests

• A practical approach for computing average temperatures at several positions 

along the main trusses (where measured temperature data were available) and a 

quantitative evaluation of the thermal model.

• A quantitative assessment of the agreement between the numerical and 

experimental results at several times into the tests.

For both tests, the modeled temperatures at the lower chord level were higher than the 

measured temperatures for most locations suggesting that error is systematic; however, at 
35

both the top chord and mid-web level the deviation of measured temperatures from the 

ANSYS temperatures were not systematic. This indicates that the concrete slab (along 

with its variability of thermal properties due to time varying moisture levels) possibly 

plays a strong role in influencing the thermal behavior of steel members near its 

proximity. Note that the effects of moisture on concrete thermophysical properties and 

moisture migration in concrete were not included in the model.

The apparent dislodgement of SFRM at the top chord was incorporated in the ANSYS 

heat transfer model. By varying the length and depth of SFRM that was dislodged at 

specific locations, an agreement between the numerical and measured temperatures was 

achieved at four locations on the upper chord along a main truss. Such analysis can 

provide approximate information about the effect of dislodgement of SFRM on steel 

temperatures and a quantitative estimate of the extent of SFRM dislodgement may be 

possible by comparing modeled and measured temperatures using a trial and error 

approach or conducting an optimization procedure with a proper objective function.

The variability of the recorded steel temperatures highlights the need for characterization 

of uncertainties in predicting temperature histories of structural elements exposed to a 

damaging fire. Sources of variability may be:

• Moisture level in the concrete slab

• Applied thickness of SFRM on steel truss elements

• Dislodgement of SFRM (extent and thickness)

• Presence (amount) of overspray

• Non-uniform heating as a result of variation in view factors as well as the 

effective radiative temperature and volume emittance. 

The time-temperatures data predicted by the numerical model are used in a structural 

model to study the effects of fire on the mechanical behavior of composite floor systems. 

The numerical heat transfer model needs to be improved by incorporating the most 

accurate data for both geometric changes in fire and thermophysical properties and heat 

transfer parameters such that predicted results are within 10  ̊C (approximately) of 

measured temperatures especially at temperatures exceeding 400  ̊C . This is because 

mechanical properties of steel (e.g., Young’s modulus, yield strength etc.) decrease 

sharply when temperatures exceed 400  ̊C. Accurate predictions of the thermal field will 

enhance the fidelity of structural predictions.

Sensitivity studies need to be conducted to determine the most important factors 

contributing to the variability in recorded temperatures. Determining uncertainties in 

time-dependent temperature histories of structural members in fire is necessary for 

predicting structural behavior in fire with a certain level of confidence. 
36

7 References

1. ANSYS Mechanical Release 12.1 (2011), ANSYS Inc., Southpointe, 275 

Technology Drive, Canonsburg, PA.

2. NIST (National Institute of Standards and Technology) (2005), Federal Building 

and Fire Safety Investigation of the World Trade Center Disaster: Final Report on 

the Collapse of the World Trade Center Towers, NIST NCSTAR 1-6B,

Gaithersburg, MD.

3. NIST (National Institute of Standards and Technology) (2005), Federal Building 

and Fire Safety Investigation of the World Trade Center Disaster: Final Report on 

the Collapse of the World Trade Center Towers, NIST NCSTAR 1-6,

Gaithersburg, MD. 

4. ASTM (1973), “Standard Test Methods for Fire Tests of Building Construction 

and Materials”, ASTM E119-73, ASTM International, Conshohocken, PA.

5. L.G. Blevins and W.M. Pitts (1999), “Modeling of Bare and Aspirated 

Thermocouples in Compartment Fire”, Fire Safety Journal. 33. 239-259.

6. S. Lamont, A.S. Usmani, and D.D. Drysdale (2011), “Heat Transfer Analysis of 

the Composite Slab in the Cardington Frame Fire Tests”, Fire Safety Journal 36: 

815-839.

7. NIST (National Institute of Standards and Technology) (2005), Federal Building 

and Fire Safety Investigation of the World Trade Center Disaster: Final Report on 

the Collapse of the World Trade Center Towers, NIST NCSTAR 1-6A,

Gaithersburg, MD.

8. ECI (1994), Eurocode 1: Basis of Design and Design Actions on Structures, Part 

2-2: Action on Structures Exposed to Fire, ENV 1991-2-2, European Committee 

for Standardization, Brussels, Belgium.

9. (a) EC1 (2002), Eurocode 1: Actions on Structures. ENV 1991, Part 1-2: General 

Actions - Actions on Structures Exposed to Fire, Brussels, European Committee 

for Standardization, Brussels, Belgium.

(b) EC2 (2002), Eurocode 2: Design of Concrete Structures. ENV 1992, Part 1-2: 

General Rules - Structural Fire Design, European Committee for Standardization, 

Brussels, Belgium.

(c) EC3 (2002), Eurocode 3: Design of Steel Structures. ENV 1993, Part 1-2: 

General Rules- Structural Fire Design, European Committee for Standardization, 

Brussels, Belgium.
37

8 Acknowledgment

The authors gratefully appreciate many helpful comments provided by Drs. 

T.P. McAllister, J.A. Main, W. Grosshandler, and F. Sadek.
38

Appendix A: Thermophysical properties

1. Steel

The thermophysical properties listed here have been taken from Ref. [7].

A. Density = 7856.2 kg/m3

B. Specific heat

Temperature Specific heat

0

C J/kg/K

27 435.3

52 452.1

77 467.2

102 481.1

127 493.9

152 505.8

177 516.4

202 526.7

227 536.3

252 545.5

277 554.8

302 563.8

327 573.2

352 582.8

377 592.9

402 603.7

427 615.5

452 628.4

477 642.3

502 657.7

527 674.5

552 692.9

577 713.6

602 736.5

627 761.2

652 788.4

677 818.1

702 850.8
39

C. Thermal conductivity

Thermal

Temperature Conductivity

0

C W/m/K

27 47.6

52 47.8

77 47.8

102 47.6

127 47.2

152 46.8

177 46.2

202 45.5

227 44.8

252 44.1

277 43.3

302 42.5

327 41.7

352 40.8

377 40.0

402 39.2

427 38.3

452 37.5

477 36.7

502 35.9

527 35.1

552 34.3

577 33.5

602 32.8

627 32.0

652 31.3

677 30.6

702 30.0

25

30

35

40

45

50

0 100 200 300 400 500 600 700 800

Thermal Conductivity, W/m/K

Temperature, 0C
40

2. Concrete

A. Density

Temperature Density

0

C kg/m3

23 2101.7

600 1951.5

800 1701.3

1000 1701.3

B. Specific heat

Temperature Specific heat

0

C J/kg/K

23 900.6

100 1000.6

800 1000.6

810 1501.0

1000 1501.0

800

1000

1200

1400

1600

0 200 400 600 800 1000 1200

Specific heat, J/kg/K

Temperature, 0C
41

C. Thermal conductivity

Thermal

Temperature Conductivity

0

C W/m/K

23 1.5

800 1.0

1000 1.0

3. SFRM

A. Density

Temperature Density

0

C kg/m3

25 237.0

50 236.3

100 230.3

200 224.8

300 222.3

400 220.5

500 219.2

600 218.4

800 361.4

1000 376.1

1200 432.4

0

0.5

1

1.5

2

0 200 400 600 800 1000 1200

Thermal conductivity, W/m/K

Temperature, 0C

100

200

300

400

500

0 200 400 600 800 1000 1200 1400

Density, kg/m3

Temperature, 0C
42

B. Specific heat

Temperature Specific heat

0

C J/kg/K

25 826.9

50 942.0

100 724.3

200 897.6

300 1020.7

400 1071.4

500 1098.2

600 1190.3

800 1259.4

1000 1326.0

1200 1392.5

C. Thermal conductivity

Thermal

Temperature Conductivity

0

C W/m/K

25 0.0

50 0.1

100 0.1

200 0.1

300 0.1

400 0.1

500 0.2

600 0.2

800 0.3

1000 0.5

1200 0.5

0

0.1

0.2

0.3

0.4

0.5

0.6

0 200 400 600 800 1000 1200 1400

Thermal conductivity, W/m/K

Temperature, 0C
43

Appendix B: Finite Element Model of the Composite Floor 

System

Figure B1: (a) side, (b) isometric, and (c) end-view of the FE model of the composite 

floor system.

(a)

(c)

(b)
44

Figure B2: Discretization in the FEM model of the composite floor system: (a) upper 

chord and slab; (b) lower chord.

(a)

(b)
45

Appendix C: Log of Observations in Test 1

Source: Reference [2]
46

Appendix D: Log of Observations in Test 2

Source: Reference [2]