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School of Urban Development

Queensland University of Technology

Numerical Models to Simulate the Thermal

Performance of LSF Wall Panels

By

Leonardus Gunawan

BE (Civil) (Queensland University of Technology)

A Thesis Submitted to the School of Urban Development,

Queensland University of Technology in Partial

Fulfillment of the Requirements for the Degree of

MASTER of ENGINEERING

August 2011

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels i

Abstract

Fire safety of buildings has been recognised as very important by the building industry

and the community at large. Gypsum plasterboards are widely used to protect light

gauge steel frame (LSF) walls all over the world. Gypsum contains free and chemically

bound water in its crystal structure. Plasterboard also contains gypsum (CaSO4.2H2O)

and calcium carbonate (CaCO3). The dehydration of gypsum and the decomposition of

calcium carbonate absorb heat, and thus are able to protect LSF walls from fires.

Kolarkar and Mahendran (2008) developed an innovative composite wall panel system,

where the insulation was sandwiched between two plasterboards to improve the thermal and

structural performance of LSF wall panels under fire conditions. In order to understand the

performance of gypsum plasterboards and LSF wall panels under standard fire

conditions, many experiments were conducted in the Fire Research Laboratory of

Queensland University of Technology (Kolarkar, 2010). Fire tests were conducted on

single, double and triple layers of Type X gypsum plasterboards and load bearing LSF

wall panels under standard fire conditions.

However, suitable numerical models have not been developed to investigate the thermal

performance of LSF walls using the innovative composite panels under standard fire

conditions. Continued reliance on expensive and time consuming fire tests is not acceptable.

Therefore this research developed suitable numerical models to investigate the thermal

performance of both plasterboard assemblies and load bearing LSF wall panels. SAFIR, a

finite element program, was used to investigate the thermal performance of gypsum

plasterboard assemblies and LSF wall panels under standard fire conditions.

Appropriate values of important thermal properties were proposed for plasterboards and

insulations based on laboratory tests, literature review and comparisons of finite

element analysis results of small scale plasterboard assemblies from this research and

corresponding experimental results from Kolarkar (2010). The important thermal

properties (thermal conductivity, specific heat capacity and density) of gypsum

plasterboard and insulation materials were proposed as functions of temperature and

used in the numerical models of load bearing LSF wall panels. Using these thermal

properties, the developed finite element models were able to accurately predict the time-

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels ii

temperature profiles of plasterboard assemblies while they predicted them reasonably

well for load bearing LSF wall systems despite the many complexities that are present

in these LSF wall systems under fires.

This thesis presents the details of the finite element models of plasterboard assemblies and

load bearing LSF wall panels including those with the composite panels developed by

Kolarkar and Mahendran (2008). It examines and compares the thermal performance of

composite panels developed based on different insulating materials of varying densities and

thicknesses based on 11 small scale tests, and makes suitable recommendations for

improved fire performance of stud wall panels protected by these composite panels. It also

presents the thermal performance data of LSF wall systems and demonstrates the superior

performance of LSF wall systems using the composite panels. Using the developed finite

element of models of LSF walls, this thesis has proposed new LSF wall systems with

increased fire rating. The developed finite element models are particularly useful in

comparing the thermal performance of different wall panel systems without time

consuming and expensive fire tests.

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels iii

TABLE OF CONTENTS

Abstract i

Table of Contents iii

List of Figures viii

List of Tables xviii

List of Charts xxii

Statement of Original Authorship xxiii

Acknowledgements xxiv

Chapter 1.0 : Introduction

1.1 : General 1-1

1.2 : Research Problem 1-3

1.3 : Research Objectives and Scope 1-4

1.4 : Research Methodology and Plan 1-4

1.5 : Contents of Thesis 1-5

Chapter 2.0 : Literature Review

2.1 : General 2-1

2.2 : Cold-formed Steel Structures 2-1

2.3 : Fire Safety 2-3

2.4 : Plasterboards 2-5

2.5 : Glass Fibre Insulation 2-7

2.6 : Rock Fibre Insulation 2-7

2.7 : Composite Insulation 2-8

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels iv

2.8 : Cassette Insulation 2-12

2.9 : Effects of Localised Fires in the Numerical Analysis of 2-14

a Building Structure

2.10 : Material Properties 2-15

2.11 : Previous Thermal Modelling 2-23

2.12 : Heat Transfer Simulation 2-28

2.13 : Thermal Performance 2-29

2.14 : SCI Publication (SCI, 1993) 2-31

2.15 : Literature Review Findings 2-34

Chapter 3.0 : Experimental Study of Thermal Properties

3.1 : General 3-1

3.2 : Test Specimens 3-1

3.3 : Test Set-up and Procedure 3-4

3.4 : Analysis of Experimental Results 3-6

3.4.1 : Typical Experimental Results of Each Specimen 3-6

3.4.2 : Calculation Methods 3-8

3.4.3 : Results for Plasterboard 3-12

3.4.4 : Results for Rock Fibre Insulation 3-16

3.4.5 : Results for Glass Fibre Insulation 3-18

3.5 : Idealised Thermal Properties to be used in Numerical 3-20

Models

3.6 : Summary 3-30

Chapter 4.0 : Finite Element Analyses of Small Scale Plasterboard Panels

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels v

4.1 : General 4-1

4.2 : SAFIR 4-1

4.3 : Limitations of SAFIR 4-3

4.3.1 : Moisture Movement 4-4

4.3.2 : Ablation 4-4

4.3.3 : Shrinkage 4-4

4.4 : GID Pre and Post Processor 4-5

4.4.1 : SAFIR Problem Types 4-5

4.4.2 : Model Geometry 4-5

4.4.3 : Materials 4-6

4.4.4 : Boundary Conditions 4-7

4.4.5 : Meshing 4-8

4.4.6 : General Data 4-9

4.4.7 : Post Processing 4-10

4.5 : Model Configuration 4-11

4.6 : Small Scale Test Specimen 1 4-12







Numerical Models to Simulate the Thermal Performance of LSF Wall Panels vi





4.17 : Summary 4-43

Chapter 5.0 : Finite Element Analyses of Load Bearing Wall Panels

5.1 : General 5-1

5.2 : Test Configuration 5-1

5.3 : Finite Element Models of Load Bearing Walls 5-6

5.4 : Load Bearing Wall Test Specimen 1 5-8

5.4.1 : Plasterboards 5-9

5.4.2 : Studs 5-12



5.5.2 : Studs 5-26



5.6.2 : Studs 5-39



5.7.2 : Studs 5-53


Numerical Models to Simulate the Thermal Performance of LSF Wall Panels vii


5.8.2 : Studs 5-66



5.9.2 : Studs 5-79

5.10 : Summary 5-89

5.11 : Improving Composite LSF Wall Panel 5-94

Chapter 6.0 : Conclusions and Recommendations

6.1 : General 6-1

6.2 : Conclusions 6-1

6.3 : Recommendations for Future Research 6-4

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels viii

List of Figures

Chapter 1.0 : Introduction

Figure 1.1 : Cold – formed Steel Structures 1-1

Figure 1.2 : LSF Plasterboard Walls 1-2


Figure 2.1 : Steel Structures 2-2

Figure 2.2 : Cold-formed Steel Cross-sections 2-2

Figure 2.3 : Fire Damaged of the Interstate Bank Building 2-3

Figure 2.4 : Sketch of Example Gypsum-panel / Steel-stud Wall 2-5

System Designs

Figure 2.5 : Glass Fibre 2-7

Figure 2.6 : Rock Fibre 2-8

Figure 2.7 : Composite Insulation Panel 2-9

Figure 2.8 : Cassette Section Stringer System 2-12

Figure 2.9 : Profiles of Cassette Systems with Interior Insulation 2-12

Figure 2.10 : Comparison of the Cold Surface Temperatures between 2-13

Different Cassette Section Systems

Figure 2.11 : Comparison of the Steel Temperatures near the unexposed 2-14

Surface

Figure 2.12 : Sketch of the Idealized Geometry of the Gypsum-panel / 2-27

Steel-stud Wall System

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels ix

Figure 2.13 : Typical Temperature History of a Steel Stud within an 2-31

LSF Wall

Figure 2.14 : Time-Temperature Curves for Glass Fibre Composite 2-31

Panel


Figure 3.1 : QUT Grinding Machine 3-2

Figure 3.2 : Plasterboard in Powder Form 3-3

Figure 3.3 : Glass Fibre Insulation 3-3

Figure 3.4 : Rock Fibre Insulation 3-3

Figure 3.5 : Al2O3 Powder 3-4

Figure 3.6 : SETARAM TGA DSC 3-4

Figure 3.7 : Aluminium Crucible Configurations 3-5

Figure 3.8 : Water and Nitrogen Control Knobs 3-5

Figure 3.9 : Typical DSC Results of Heat Flow versus Time for 3-6

Gypsum Plasterboard

Figure 3.10 : Typical DSC Results of Heat Flow versus Time for Rock 3-7

Fibre

Figure 3.11 : Typical DSC Results of Heat Flow versus Time for Glass 3-7

Fibre

Figure 3.12 : Typical Continuous Cp without the DSC Results of 3-8

Reference Material

Figure 3.13 : Specific Heat of Al2O3 from ASTM E 1269 (ASTM, 2005) 3-9

Figure 3.14 : Typical Continuous Cp (mass) with the DSC Results of 3-10

Reference Material

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels x

Figure 3.15 : Typical Continuous Cp (volume) with the DSC Results of 3-11

Reference Material

Figure 3.16 : Heat Flow versus Time for Plasterboards 3-13 – 3-14

Figure 3.17 : Mass Loss in Plasterboards 3-15

Figure 3.18 : Specific Heat of Plasterboards 3-15

Figure 3.19 : Rock Fibre After DSC Test 3-16

Figure 3.20 : Heat Flow versus Time for Rock Fibre Insulation 3-16

Figure 3.21 : Mass Loss in Rock Fibre Insulation 3-17

Figure 3.22 : Specific Heat of Rock Fibre Insulation 3-17

Figure 3.23 : Glass Fibre after DSC Test 3-18

Figure 3.24 : Heat Flow versus Time for Glass Fibre Insulation 3-18

Figure 3.25 : Mass Loss in Glass Fibre Insulation 3-19

Figure 3.26 : Specific Heat of Glass Fibre Insulation 3-19

Figure 3.27 : Specific Heat of Plasterboard Reported by Various 3-23

Researchers

Figure 3.28 : Proposed Specific Heat of Plasterboard 3-23

Figure 3.29 : Proposed Thermal Conductivity of Plasterboard 3-25

Figure 3.30 : Proposed Relative Density of Plasterboard 3-26

Figure 3.31 : Specific Heat of Rock Fibre Insulation 3-27

Figure 3.32 : Thermal Conductivity of Rock Fibre Insulation 3-27

Figure 3.33 : Specific Heat of Glass Fibre Insulation 3-28

Figure 3.34 : Thermal Conductivity of Glass Fibre Insulation 3-29

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xi


Figure 4.1 : SAFIR Problem Types 4-5

Figure 4.2 : Typical GID Geometry 4-6

Figure 4.3 : GID Interface for Material Condition 4-6

Figure 4.4 : Small Scale Test Specimen 6 with Material 4-7

Figure 4.5 : Specimen 3 Boundary Conditions 4-7

Figure 4.6 : Mesh Generation Dialog Box 4-8

Figure 4.7 : Summary of Mesh Generated 4-8

Figure 4.8 : Generated Finite Element Mesh 4-9

Figure 4.9 : SAFIR Problem Data 4-9

Figure 4.10 : GID Post-process Interface with Temperature Contours 4-10

Active

Figure 4.11 : Test Set-up of Gypsum Plasterboard (Kolarkar, 2010) 4-11

Figure 4.12 : Small Scale Test Specimen 1 4-13

Figure 4.13 : Time-temperature Profiles of Test Specimen 1 (13 mm 4-13

Plasterboard) from Experiment and FEA

Figure 4.14 : Specimen 1 Temperature Distributions from FEA 4-14


Figure 4.16 : Time-temperature Profiles of Test Specimen 2 (16 mm 4-17

Plasterboard) from Experiment and FEA

Figure 4.17 : Specimen 2 Temperatures Distribution from FEA 4-17 – 4-18


Figure 4.19 : Time-temperature Profiles of Test Specimen 3 4-20

(13 & 16 mm Plasterboards) from Experiment and FEA

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xii

Figure 4.20 : Specimen 3 Temperature Distributions from FEA 4-21


Figure 4.22 : Time-temperature Profiles of Test Specimen 4 (Two 4-24

16 mm Plasterboards) from Experiment and FEA

Figure 4.23 : Specimen 4 Temperature Distributions from FEA 4-24 – 4-25

Figure 4.24 : Time-temperature Profiles of Test Specimen 5 (Three 4-26

16 mm Plasterboards) from Experiment and FEA



Figure 4.27 : Time-temperature Profiles of Test Specimen 6 from 4-29

Experiment and FEA



Experiment and FEA




Experiment and FEA



Experiment and FEA




Experiment and FEA

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xiii



Experiment and FEA


Figure 4.41 : Gypsum Plasterboard Failure Time vs. Thickness 4-44

Figure 4.42 : Ambient Side Time-temperatures Profiles for All 4-45

Specimens


Figure 5.1 : LSF Wall Panel 5-3

Figure 5.2 : Gas Furnace 5-4

Figure 5.3 : Thermocouple Locations for Load Bearing Wall 5-4 – 5-5

Specimens

Figure 5.4 : Loading Frame Arrangement (Kolarkar, 2010) 5-5 – 5-6

Figure 5.5 : Complete Set-up of Load Bearing Wall Test 5-6

Figure 5.6 : Load Bearing Wall Finite Element Mesh 5-7

Figure 5.7 : Time-temperature Profiles of Specimen 1 from 5-9 – 5-11

FEA & Experiment

Figure 5.8 : Time-temperature Profiles of Stud 1 from FEA 5-13 – 5-14

& Experiment


& Experiment


& Experiment


& Experiment

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xiv

Figure 5.12 : Average Time-temperature Profiles of Studs 1 to 4 5-7

from FEA and Experiment

Figure 5.13 : Temperature Distributions from FEA for Test Specimen 1 5-21


FEA & Experiment


& Experiment


& Experiment


& Experiment


& Experiment




Figure 5.21 : Construction of Test Specimen 3 5-35


FEA & Experiment


& Experiment


& Experiment


& Experiment

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xv


& Experiment






FEA & Experiment


& Experiment


& Experiment


& Experiment


& Experiment






FEA & Experiment


& Experiment


& Experiment

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xvi


& Experiment


& Experiment






FEA & Experiment


& Experiment


& Experiment


& Experiment


& Experiment




Figure 5.53 : Time-temperature Profiles from FEA for Steel Studs 5-92 – 5-93

Figure 5.54 : Time-temperature Profiles from FEA for Ambient 5-94

Side Gypsum Plasterboards

Figure 5.55 : FEA Experiment 1 (FEA1) 5-95

Figure 5.56 : Comparison between FEA1 and Test Specimen 6 5-97

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xvii

Figure 5.57 : Time-temperature Profiles of FEA1 5-99

Figure 5.58 : FEA Experiment 2 (FEA2) 5-99

Figure 5.59 : Time-temperature Profiles of Steel Studs 5-101

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xviii

List of Tables


Table 2.1 : Thermal Conductivity of Rock Fibre Insulation 2-8

Table 2.2 : Details of LSF Wall Specimens Tested by Gunalan (2009) 2-10

Table 2.3 : Details of LSF Wall Specimens Tested by Kolarkar (2010) 2-11

Table 2.4 : Apparent Thermal Properties of Firecode C Core Type X 2-16

Gypsum Board

Table 2.5 : Apparent Thermal Properties of Insulation Materials 2-17

Table 2.6 : Summary of FRM Fire Resistance Tests on Load Bearing 2-30

LSF Walls

Table 2.7 : Fire Resistance of Typical Floors, Walls and Partitions 2-32

Comprising Cold-formed Steel Sections and Planar Board

Protection

Table 2.8 : Strength Reduction Factors for Cold-formed Steel at 2-33

Elevated Temperatures

Table 2.9 : Limiting Temperatures of Beams and Columns using 2-34

Cold-formed Steel Sections


Table 3.1 : Initial Mass of Materials Used in the DSC Test 3-2

Table 3.2 : Al2O3 Used in the DSC Test 3-4

Table 3.3 : Proposed Specific Heat of Plasterboard 3-24

Table 3.4 : Proposed Thermal Conductivity of Plasterboard 3-25

Table 3.5 : Proposed Relative Density of Plasterboard 3-26

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xix

Table 3.6 : Proposed Thermal Conductivity of Rock Fibre Insulation 3-28

Table 3.7 : Proposed Thermal Conductivity of Glass Fibre Insulation 3-29


Table 4.1 : Details of Plasterboard Test Specimens (Kolarkar, 2010) 4-12

Table 4.2 : Comparison of Experimental and Finite Element Analysis 4-15

Results for Test Specimen 1





















Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xx

Table 4.13 : Summary of Gypsum Plasterboard Small Scale Test 4-43

Table 4.14 : Summary of Glass and Rock Fibre Small Scale Test 4-45


Table 5.1 : Load Bearing Wall Configurations in Kolarkar’s 5-2

(2010) Fire Tests

Table 5.2 : Meshing Details 5-7

Table 5.3 : Comparison of Finite Element Analysis and 5-11 – 5-12

Experimental Results of Plasterboards for Test

Specimen 1

Table 5.4 : Comparison of Finite Element Analysis and 5-20

Experimental Results of Steel Studs for Test

Specimen 1



Specimen 2



Specimen 2



Specimen 3



Specimen 3



Specimen 4

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xxi



Specimen 4



Specimen 5



Specimen 5



Specimen 6



Specimen 6


Experimental Results of Load Bearing Wall Tests

Table 5.16 : Failure Temperature of Steel Studs and Ambient Side 5-96

Plasterboards from FEA

Table 5.17 : Failure Temperature of Steel Studs and Ambient Side 5-96

Plasterboards from Experiment (Kolarkar, 2010)

Table 5.18 : Failure Temperature of FEA1 5-97

Table 5.19 : Failure Temperature of FEA1 & FEA2 5-100

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xxii

List of Charts


Chart 3.1 : Process to Determine the Idealised Thermal Properties 3-21

of Gypsum Plasterboard

Chart 3.2 : Process to Determine the Idealised Thermal Properties 3-22

of Rock and Glass Fibre Insulations

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xxiii

Statement of Original Authorship

The work contained in this thesis has not been previously submitted to meet

requirements for an award at this or any other higher education institution. To the best

of my knowledge and belief, the thesis contains no material previously published or

written by another person except where due reference is made.

Leonardus Gunawan

Signed :____________________________________________________

Date :____________________________________________________

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels xxiv

Acknowledgements

The research described in this report was carried out in the School of Urban

Development, Queensland University of Technology, Australia.

The author would like to thank his supervisor Prof. Mahen Mahendran of the

Queensland University of Technology for his inspiration, guidance and enthusiasm.

Thanks also to my fellow researchers Poologanathan Keerthan, S. Gunalan, B.

Baleshan, and Anthony Deloge Ariyanayagam for their help during my research work

and to the technical staff in the laboratory, Eric Martinez for his help in DSC test.

Thanks also to Queensland University of Technology for providing the necessary

facilities and support to conduct this research project.

The author would also like to thank his parents, fiancé and brother for their immense

support to the completion of this thesis.

Introduction

Numerical Models to Simulate the Thermal Performance of LSF Wall Panels 1-1

Chapter 1

Introduction

1.1. General

Light gauge cold-formed steel frame structures are increasingly used in commercial and

residential buildings because of their non-combustibility, dimensional stability, ease of

installation and other good features (Figure 1.1). The growth of Light Gauge Steel

Frame (LSF) systems is expected to increase the demand for economical solutions,

where specific performance is required, such as in the area of fire resistance. Achieving

sufficient fire resistance is to prevent or delay the spread of fire and to ensure building

integrity. With increasing use of LSF systems in load bearing applications, the demand

for LSF systems with improved fire resistance ratings has increased. There is a need to

develop new LSF wall and floor systems with increased fire resistance rating to replace

the conventional LSF systems with plasterboard protection and cavity insulation.

Innovative fire protection systems are therefore essential without simply adding on

more gypsum plasterboards, which is inefficient.

Under fire conditions, cold-formed thin-walled steel sections heat up quickly resulting

in fast reduction in their strength and stiffness. Light gauge cold-formed steel joist or

stud sections are commonly used in planar structural floor and wall systems with

plasterboards on both sides as fire protection (Figure 1.2). This provides protection to

steel joists and studs during building fires, delaying the temperature rise in the cavity.

Figure 1.1: Cold- formed Steel Structures (Al Maher Contracting, 2011)

Introduction


Figure 1.2: LSF Plasterboard Walls (Isolgomma, 2011)

Fire safety design is an essential component of building design. A properly designed

building system greatly reduces the hazards to life and limits property loss. The research

on fire safety design commenced many years ago and the resulting development of

sound fire engineering principles has brought significant reduction to the cost of fire

protection. However, the traditional method of using many layers of fire protection

material is still continuing although it is approximate and conservative.

In Australia, there has not been any research done in this area except for the recent work

at QUT by Kolarkar and Mahendran (2008). They developed a new composite panel

system in which insulation was used externally between plasterboards instead of the

conventional cavity insulation located within the stud/joist space and investigated its

application for wall systems. They carried out full scale tests to investigate the

effectiveness of stud walls protected by the new composite systems. Effect of cavity

insulation and its location within the wall system was also investigated. Their tests

demonstrated that the use of this new composite panel system improved the fire rating

of stud wall systems. But its increased fire rating could not be determined using the

currently available design methods.

Nowadays, computational techniques are becoming increasingly important as research

tools in structural fire protection because they provide engineers with a better

understanding and interpretation of experimental results. These techniques can also be

Introduction


used to determine the increased fire rating for the new composite panel system that has

been developed by Kolarkar and Mahendran (2008).

1.2. Research Problem

The research on LSF stud wall systems under fire conditions is relatively recent and the

behaviour of wall insulation, plasterboards, and other components in the LSF wall

systems is not fully understood. The relationships between the fire resistance rating and

the various parameters such as the number of plasterboards, thickness of the

plasterboard, grade and thickness of steel and cross sectional shape of joists and screw

spacing, and load ratio, are not well understood. Despite this, the LSF wall design

continues to be based on time consuming and expensive full scale fire tests.

Nowadays fire protection rating is increased by simply adding on more plasterboards.

This conventional system is inefficient since it uses more materials thus making the wall

and floor systems thicker and heavier without a significant increase in their fire rating

performance. Kolarkar and Mahendran (2008) have recently developed a new

composite panel system with increased fire rating as shown by their full scale fire tests.

However, there is limited understanding of both the thermal and structural performances

of this LSF wall system using the new composite panels.

There is an immediate need for the development and use of suitable numerical and/or

analytical models of conventional and new LSF wall systems to improve the

understanding of their thermal and structural performances beyond what was gained

from the limited full scale fire tests. To address this problem QUT researchers have

recently commenced their research into the structural performance of LSF wall systems

using finite element models. However, no research has been undertaken on thermal

modeling of LSF wall systems, in particular the new system developed by Kolarkar and

Mahendran (2008). Therefore this research will concentrate on developing suitable

finite element models to investigate the thermal behavior of LSF wall systems and to

improve the knowledge and understanding of their thermal performance while

developing the required time-temperature profiles for these wall systems under standard

fire conditions. These time-temperature profile results can then be successfully used in

their structural modeling phases.

Introduction


The use of accurate thermal properties in finite element modeling is essential if valid

thermal performance simulations are needed. Hence this research also undertook a task

to develop accurate thermal properties of LSF wall materials using appropriate

experiments and available literature.

1.3. Research Objectives and Scope

The main objective of this study is to develop suitable finite element models to simulate

the thermal behaviour of plasterboard assemblies and load bearing LSF wall systems

under fire conditions and use them in a study to improve the understanding of the effect

of relevant parameters on their thermal performance. The scope of this research project

is defined by the following specific objectives:

1. To measure the thermal properties such as specific heat and mass loss of gypsum

plasterboard, glass and rock fibre insulations tested by Kolarkar (2010).

2. To develop idealised thermal properties of gypsum plasterboard, glass and rock

fibre insulations to be used in the finite element models of LSF wall panels.

3. To develop finite element models capable of simulating the thermal behaviour of

plasterboard assemblies and load bearing LSF wall panel systems including the

new LSF wall system with a composite panel under fire conditions with an

acceptable accuracy and validate them using the available experimental results

provided by Kolarkar (2010).

4. To study the thermal performance of plasterboard assemblies and load bearing

LSF wall panels under fire conditions

1.4. Research Methodology and Plan

Phase 1: Literature Review

Independent reading was undertaken to gain essential background knowledge regarding

the thermal performance of LSF wall systems under fire conditions. Following areas

were the focus of the literature review.

Introduction


Thermal and structural behaviour and failure mechanisms of the new composite

LSF wall panel.

Heat transfer through different types of insulation (void, steel, glass fibre, rock

fibre).

Theory and application of Finite Element Method (FEM) especially using

SAFIR and GID.

Phase 2: Measurements of Thermal Properties

This phase considered the influence of various thermal properties such as specific heat,

thermal conductivity and mass loss. These properties were measured using a DSC

machine and standard measurement procedures in ASTM E 1269 (ASTM, 2005).

Phase 3: Development of Numerical Model and Validation

Finite element models simulating the thermal behaviour of LSF wall systems were

developed using a finite element software called SAFIR and GID. This research focused

on thermal modelling of small scale plasterboard panels and load bearing LSF walls

tested by Kolarkar (2010). The accuracy of the developed finite element models was

validated using the results from six load bearing walls and eleven small scale tests.

1.5. Contents of Thesis

Chapter 1 : Gives a general introduction to cold-formed steel stud wall systems and

the objectives of the current research.

Chapter 2 : Past research on the thermal performance and modelling of steel stud

wall systems at elevated temperatures.

Chapter 3 : Presents the details of DSC tests to measure specific heat and heat loss

of gypsum plasterboard, glass and rock fibre insulations. It also presents the idealized

thermal properties of these materials used in LSF wall assemblies.

Introduction


Chapter 4 : Presents the details of finite element modelling using SAFIR and GiD

and validation of the developed finite element models using the eleven small scale test

results of Kolarkar (2010).

Chapter 5 : Presents the validation of the developed finite element models using the

six load bearing wall test results of Kolarkar (2010).

Chapter 6 : Presents the main findings and recommendations.

Literature Review


Chapter 2

Literature Review

2.1. General

This literature review includes the details of cold-formed steel structures, fire safety,

LSF wall systems and different types of insulation. Further, it explores the current state

of knowledge of fire performance, fire resistance of LSF walls and thermal modelling.

Finally, the literature review summarises the current state of knowledge and the

contribution to the current research.

2.2. Cold-formed Steel Structures

Cold-formed steel products have enjoyed significant growth in recent years. They may

be utilized in various forms on commercial, industrial and residential construction

projects today (Figures 2.1 and 2.2). Their strength, light weight, versatility, non-

combustibility, and ease of production have encouraged architects, engineers, and

contractors to use cold-formed steel products which can improve structural function and

building performance, and provide aesthetic appeal at lower cost.

The reasons behind the growing popularity of these cold-formed steel products include

their ease of fabrication, high strength to weight ratio and suitability for a wide range of

applications. These advantages can result in more cost-effective designs compared with

hot-rolled steel members, especially in short-span applications.

Cold-formed steel members can be produced in a wide variety of sectional profiles, the

most commonly used of which are the C (channels) and the Z sections (Figure 2.2). The

thickness of the materials most frequently used for these structural members ranges

from about 0.4 mm to 6.4 mm.

Literature Review


Figure 2.1: Steel Structures (Top Free Biz, 2011)

Figure 2.2: Cold-Formed Steel Cross-Sections (Risto Hara, 2000)

Although these cold-formed steel members are considered to be more efficient than hot-

rolled steel members, the versatility of the different shapes and sizes of cold-formed

steel sections that are currently available allow the sections to be used effectively as

primary stud walls, floor beams, roof trusses and partitions.

In LSF stud wall construction, small size steel members are used. In a wall panel, studs

(vertical members) and tracks (horizontal members) are normally used. Studs carry the

vertical load with tracks connecting the studs to make the frame. These members are

manufactured by cold rolling of 0.5 to 1.5 mm steel strips. Hence fire resistance must be

based on protective materials, by far the most common being gypsum board. Gypsum

board has fire resistance properties better than most of the other similar materials

because of the moisture in the gypsum crystals.

The wall panels are typically constructed by first connecting sheathing boards to the

frame with self-drilling screws. Lee (1999), Telue (2001) and Tian et al. (2007) present

some of the studies on the structural (axial compressive) capacity of these wall panels at

ambient temperature. Insulation materials are added when additional fire rating is

Literature Review


required. These panels can be easily assembled to fabricate load-bearing or non-load-

bearing walls.

2.3. Fire Safety

Unwanted fire is destructive and causes many deaths and billions of dollars of property

loss each year. People around the world expect that their homes and work places will be

safe from the ravages of unwanted fire. Unfortunately fire can occur in almost any kind

of building (Figure 2.3), often when least expected. The safety of occupants depends on

many factors in the design and construction of the building. Most often verification of

the fire-resistance of light frame structures is in the time domain, where proprietary

ratings are compared with the code specified fire resistance, or with the calculated

equivalent time of a complete burnout.

Figure 2.3: Fire Damaged of the Interstate Bank Building (Los Angeles Fire

Department Historical Archive, 1988)

The failure criteria for fire safety can be in terms of Structural adequacy, Integrity or

Insulation. Integrity is defined in AS 1530.4 (SA, 1997) as the ability to resist flames or

smoke passing through the section. The three fire resistance requirements, stability,

Insulation and integrity, should be satisfied when walls are used in fire rated

construction. LSF walls lose their strength at elevated temperatures since they are made

of thin steel sheets. Therefore LSF walls are usually covered with plasterboards and

insulation materials to make a wall assembly that can withstand the required fire

exposure.

Literature Review


Assessment of integrity must be done in full-scale testing because small-scale tests

cannot assess factors such as shrinkage in large sheets of gypsum board or cracking due

to structural deformations.

Based on the ISO 834 (ISO 1999) and AS 1530.4 (SA, 1997) criteria, the assembly is

considered to have failed the test by the insulation criterion when the average

temperature rise on the unexposed surface exceeds 140°C, or the maximum temperature

rise at any point exceeds 180°C.

The structural adequacy criteria shall be deemed to have occurred when either a

collapse occurs or when the deflection exceeds its limit. The strength of LSF assemblies

is mainly in the steel members themselves and not the lining materials. Lining materials

are essential for providing lateral stability to the structural members, but their

contribution to overall strength and stiffness is small.

The fire resistance of LSF systems protected by plasterboard depends on several

important interrelated properties: The insulating capacity of the board protects the

internal structural members and delays temperature rise on unexposed surfaces; The

ability of the board to remain in place and not disintegrate or fall off after dehydration;

(the extent to which glass fibre reinforcing and closely spaced fixings can hold the

board together after the gypsum has dehydrated); Resistance to shrinkage which usually

causes cracking within the board or separation at joints between sheets (glass fibres and

additive such as vermiculite found to be controlling shrinkage); The ability of the core

material to resist ablation from the fire side during extreme fire exposure.

Sultan et al. (1998) concluded that the joist spacing did not affect the fire resistance of

assemblies. Screw spacing from the gypsum board edge, cavity insulation, and the

number of gypsum boards had a significant effect on the fire performance of wall

assemblies. However, the effect of steel grades, steel sheet thickness and joist section

properties were not considered. There were mixed outcomes in relation to structural

load and type of insulation depending on the type of joists (wood or steel).

Literature Review


2.4. Plasterboards

Figure 2.4 adopted from Sultan (2001), is a sketch of the wall system design. In general,

two arbitrary-thickness gypsum wall panels are mounted one on either side of an array

of vertical steel studs. In practice, each of the two panels shown can involve a single

thickness of gypsum board or a sandwich-type multiple-thickness design of two or more

well-contacted boards. Figure 2.4 illustrates two particular assembly designs. One of

these is referred to as 1x1-type assembly, since each of the two panels involves a single

layer of gypsum board and the other involves a two-layer construction. The studs,

spaced at regular intervals, from an unfilled air gap between the panels. Also, the studs

are typically fabricated from relatively thin steel (the studs used in the experimental

study Sultan (2001) were 0.46mm thick and they do not contribute much to the heat

transfer between the panels, the spacing of the studs is several times the thickness of the

air gap. It is only in relatively“sparse“regions of the wall system that the presence of the

studs introduces two-dimensional consideration into the wall system geometry and heat

transfer (Franssen, 1999).

Figure 2.4: Sketch of example gypsum-panel/steel-stud wall system designs

(adopted from ASTM E199, 1998; ISO-834 Switzerland, 1992)

When gypsum board lining is heated during a fire, temperature on the exposed face will

increase steadily until about 100°C is reached, at which time there will be a delay while

the water of crystallization is driven off. As the heating continues, the 100°C

temperature plateau will progress slowly through the board, until the entire board has

been dehydrated. Hence this protects the steel frame from heat.

Literature Review


After dehydration the gypsum has almost no strength because it has been converted to a

powdery form. Any residual strength depends on glass fibre reinforcing to hold the

board together. Temperatures within the board will rise steadily after dehydration is

completed, leading to increased temperatures in the cavity and in the framing members.

The strength of steel joist depends on the temperature of the joist and the level of

stability provided to the joist by the lining materials. As a fire progresses, steel framing

will lose strength due to increased temperatures, but long periods of fire resistance can

be achieved if the lining on the fire side remains in place. Resistance to fire also

depends on how much heat is transferred across the cavity and through the lining on the

unexposed side.

The insulation criterion for fire resistance requires that the temperature on exposed face

remains below a certain critical temperature, so there is no danger of ignition on the

unexposed surface and subsequent fire growth. From the past study using insulation as

cavity makes the failure of joists to happen earlier because of thermal bowing of steel

joists (Sultan 1998).The new composite system of Kolarkar and Mahendran (2008).

(Using insulation between plasterboards externally) gave better results than using

insulation in the cavity.

In addition to this, other important factors affecting the fire resistance are the thickness

of the gypsum board, the quality of the board material, the details of the construction

and the fixings and standard of workmanship. Recent tests show a significant increase

in fire resistance if the screws are located at least 35 mm from the edge of the gypsum

board (Sultan, 1998).

Multiple layers of thin gypsum boards may be cheaper and lighter to fix than one thick

board, but multiple layers do not usually provide the same fire resistance as a single

layer of the same total thickness, because the outer layers can fall off sequentially,

leading to much greater thermal exposure to the inner board. An advantage of multiple

boards is that the joints between the sheets can be staggered, reducing the likelihood of

early flame penetration into the cavity, especially if sheet joints are not on studs. If more

than one layer is used, the inner layer is not usually taped or stopped at joints.

Literature Review


2.5. Glass Fibre Insulation

Glass fibre insulation (see Figure 2.5) plays a significant role in energy conservation in

buildings. These insulators are manufactured in the form of thick rectangular sheets and

are fitted in the walls. Physical experiments have shown that as the porosity of this

insulating material increases, i.e. the amount of glass fibre decreases, the effective heat

transfer coefficient of the material also decreases. This is in accordance with the fact

that the coefficient of thermal conductivity of glass is quite large compared to that of

air. But this phenomenon continues only up to a certain level of porosity. If the porosity

crosses that level, the heat transfer coefficient of the insulating material again starts to

rise with porosity. It is indicating that the quality of the glass fibre insulators can be

improved by reducing the non-homogenous in the glass fibre density as much as

possible (Sundar et al, 2006).

Figure 2.5: Glass Fibre

2.6. Rock Fibre Insulation

Rock fibre insulation conducts heat very well, but when pressed into rolls and sheets

their ability to partition air makes them excellent heat and sound insulators (Figure 2.6).

Rock fibre is made from natural minerals like basalt or diabase. In addition to providing

thermal insulation, it also absorbs sound and, with a vapour retarder, helps control

condensation. Because they are non-combustible and have melting temperature in

excess of 1000°C, they are also used to prevent the spread of fire. Although not immune

to the effects of a sufficiently hot fire, the fire resistance of glass fibre, rock fibre and

Literature Review


ceramic fibres makes them common building materials when passive fire protection is

required, being used as spray fireproofing, in stud cavities in drywall assemblies and as

packing materials in fire stops. Whilst rock fibre‟s greater density (greater than

100kg/m3 according to ASTM C612-93) provides unique benefits, it also means that it

cannot be compressed. Thermal conductivity of insulation is tested in accordance with

German standards (DIN 52612) and used by Australasian Insulation Supplies.

Table 2.1: Thermal Conductivity of Rock Fibre Insulation (Alfawakhiri, 2001)

Temperature(°C) 50 100 150 200 250 300

Thermal conductivity

(W/m °C) 0.038 0.043 0.049 0.058 0.067 0.078

Rock fibre specific heat is 0.84 kJ/kg.°C.

Figure 2.6: Rock Fibre

2.7. Composite Insulation

Past research has produced contradicting results about the benefits of cavity insulation

to the fire rating of stud wall systems. Because of the very low conductivity of the

cavity insulating material as compared to steel, most of the heat gets directed along and

across the steel studs while cavity insulation acts as the heat sink thus raising their

temperatures much faster in the case on non-cavity insulated specimens. This makes the

presence of cavity insulation a threat to the survival of steel frame under fire conditions.

Literature Review


In many tests, it was concluded that the cavity insulation reduces the fire resistance of

LSF wall panels and there was a need to develop new wall systems with increased fire

rating. This resulted in the new development of a new composite panel by Kolarkar and

Mahendran (2008). The new stud wall systems was built with the insulation sandwiched

between the plasterboards on either side of the steel wall frame instead of being placed

in the cavity as shown in Figure 2.7.

Figure 2.7: Composite Insulation Panel (Kolarkar, 2010)

Externally insulated wall panels such as the new composite panel can offer a much

higher level of protection to the studs as they are installed on the fire side of the studs

thus minimizing the transfer of heat by radiation and conduction. Hence the quality of

insulation used externally would directly influence the level of fire protection offered to

the studs.

Recently Kolarkar and Mahendran (2008) undertook research into the fire resistance of

LSF stud wall panels based on nine full scale tests of load bearing walls and nine small

scale tests of non-load bearing walls. The test frames were made of 1.15mm G500 cold-

formed steels whereas the plasterboard used had a nominal thickness of 16 mm. Glass

fibres, rock wool and cellulosic fibre were used as the insulation materials. The

traditional system of putting the insulation inside the cavity was found to be inefficient

and innovative composite panel was introduced to increase the fire resistance of wall

panels. The idea was to use the insulation layer outside the steel frame as shown in

Figure 2.7. Test results showed that the new composite panel improved the fire rating of

LSF wall panels.

Table 2.3 shows the details and the failure times of test specimens tested by Kolarkar

(2009). The first specimen was tested in ambient condition whereas all other specimens

were tested under standard fire conditions. These specimens were tested under a

gypsum board

Rock fibre

Steel

Literature Review


constant load of 15 kN/Stud, ie. Load ratio of 0.2. These results demonstrated the

improved fire performance of LSF wall assemblies when insulation was used externally

between plasterboards instead of using in the cavity. However, Kolarkar‟s study was

limited to experimental work and further numerical modelling of the new stud walls is

required to investigate the possibility of improving the new composite system further.

Hence there is a need to investigate both the non-load bearing and load-bearing walls

made of the new composite panel to fully understand their structural and thermal

behaviour and to improve their fire resistance rating. Recently Gunalan (2009) has

commenced his research into the structural behaviour of LSF walls at QUT. He

conducted three full scale fire tests of LSF walls, but under a higher load ratio of 0.4.

Table 2.2 shows the details and results of Gunalan‟s tests.

Table 2.2: Details of LSF Wall Specimens Tested by Gunalan (2009)

Test

No.

Configuration Load

Ratio

External

Insulation

Failure Mode Failure

Time

(min)

01 0.2 Glass Fibre Structural 118

02 0.4 Glass Fibre Structural 108

03 0.4 Rock Fibre Structural 134

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Table 2.3: Details of LSF Wall Specimens Tested by Kolarkar (2010)

Test

Sp. Configuration Condition Insulation

Failure Time

(minutes)

01 (P) Ambient None -

02 (P) Fire None 53

03 (P) Fire None 111

04 (P) Fire

Glass Fibre

101

(Cavity Insulation)

05 (P) Fire

Rock Fibre

107

(Cavity Insulation)

06 (P) Fire

Cellulose Fibre

110

(Cavity Insulation)

07 (P) Fire

Glass Fibre 181

(Unexpected

(External Insulation) furnace failure)

08 (P) Fire

Rock Fibre

136

(External Insulation)

09 (P) Fire

Cellulose Fibre

124

(External Insulation)

Literature Review


2.8. Cassette Insulation

Cassette sections are one of the recent applications of cold-formed thin-walled steel in

wall panels. Figure 2.8 shows the dimensions of a typical cassette section.

Figure 2.8: Cassette Section Stringer System (Feng et al, 2003)

The flanges of a cassette section would occupy the space of the web of conventional

steel channel section in a steel stud panel. In practical construction, the wide web of

cassette sections are joined together to form a continuous surface. The intention is to

utilise diaphragm action in the web plane of the cassette section to eliminate bracing in

conventional steel stud construction. Although there is an increased use of steel, the

construction speed is faster, resulting in an overall reduction cost. The fire performance

of a cassette system may be influenced by the double thickness at the junction of two

cassettes and the continuous web surface of the cassette section.

System 1 System 2 System 3 System 4

Figure 2.9: Profiles of Cassette Systems with Interior Insulation (Feng et al, 2003)

Literature Review


Feng et.al (2003) investigated the thermal performance of this type of wall panel (Figure

2.9). Since the continuous steel sheet on the fire exposed side attracts more heat, the

unexposed surface and steel temperatures in Cases 1 and 2 are much higher than those

in Cases 3 and 4, respectively. Furthermore, steel temperatures near the unexposed

surface (Figure 2.10 and 2.11) in Case 4 are much higher than that in Case 1 initially,

due to using one layer of gypsum board. However, due to reduced heat attracted by the

narrow webs on the exposed side, temperatures in Case 4 eventually become lower than

in Case 1. The main conclusion is that when using cassette sections, it is better to have

the continuous steel sheet on the unexposed side of the wall. This ensures that the

system only attracts a small amount of heat from fire exposure on the narrow webs. But

during real fire events it is almost impossible to predict from which direction the fire

will come from, therefore further development is needed to improve the fire

performance of cassette systems.

Figure 2.10: Comparison of the Cold Surface Temperatures between Different

Cassette Section Systems (Feng et al, 2003)

Literature Review


Figure 2.11: Comparison of the Steel Temperatures near the Unexposed Surface

(Feng et al, 2003)

2.9. Effects of Localised Fires in the Numerical Analysis of a Building Structure

When structural elements are tested against fire in a furnace, every precaution is taken

in order to have a uniform spatial distribution of the temperature in the furnace or, more

precisely, to have a uniform thermal attack on the elements. In real buildings, the fully

developed fire that usually takes place in the compartment is considered to be

represented accurately enough by a one-zone situation, which means that the conditions

in terms of gas temperatures or incident heat flux to the structure, are uniform.

This uniform situation has direct consequences on the numerical simulations that are

performed to model the behaviour of the structure. The temperature distribution in the

flat elements such as walls, floors, and ceilings is essentially one-dimensional (1D),

with a gradient only across the thickness of the slab. In linear elements, such as beams

and columns, the temperature distribution is essentially two-dimensional (2D) with no

variation along the length of the elements. This particular temperature distribution is of

course taken into account in the analyses and, in beams for example, as long as the

cross-section remains the same, the same temperature distribution is considered for

every longitudinal point of integration. The same holds in the slabs for every point of

integration in the plane of elements. The temperature distribution can have a 2D (for

slabs) or three-dimensional (3D) (for beams) pattern near the edges of the compartment

Literature Review


because of the influence of the adjacent cold compartments, but this effect is strongly

localised and is usually neglected.

At the beginning of every fire it is localised in the compartment before it turns into a

fully developed fire, this preliminary phase is usually disregarded for the analysis of the

structural behaviour because the low temperatures associated to this phase are

considered to have negligible effects on the structure. In some cases, even this localised

fire may be a threat for the structure. One example is a localised fire under a statically

determinate steel truss girder; losing the one member of the truss that is located just

above the fire leads to the loss of the whole girder. Some fires keep a localised character

during the entire duration.

The methodologies that are used for analysing the fire behaviour of a structure that is

subjected to a uniform thermal situation cannot be applied when the fire is localised.

The concept of “zoning” can be applied. The structure is divided into several zones in

which the situation is approximated as uniform.

2.10. Material Properties

Thermo physical and mechanical properties are widely reported as a function of

temperature for several structural and reinforcing steels (Lie 1972, 1992; Pettersson et

al., 1976; Pettersson 1986; Kirby and Preston 1988; Schleich 1993). In addition to these,

the more related expressions to this study are discussed and presented next.

(a). Plasterboards

Gypsum board is widely used for interior lining in domestic housing and commercial

office buildings, and is the most commonly used lining material to provide light frame

structures with fire resistance. Typical Gypsum board has a density between 550 and

850 kg/m3. Most gypsum boards are made with a thickness between 10 and 20 mm.

There are three broad types of gypsum board, usually known as Regular board, Type X

board, and Special purpose boards. Regular gypsum board, a generic product, is not

required to have any fire resistance rating and has low density with no reinforcing. Type

Literature Review


X board is also a generic product, which is used as a fire resistant material. Special

purpose boards are proprietary products made with more glass fibres and more core

additives to obtain enhanced fire or structural performance.

Thermal conductivity also depends on the density of the gypsum board. Its value above

about 400°C will be affected by the presence of shrinkage cracks in the gypsum board,

which will depend on the formulation of the individual board and the type of fire.

Sultan (1996) reported that fall-off of plasterboard occurs when the unexposed face of

the board reaches about 600°C (Buchanan and Gerlich, 1997). However, Kolarkar and

Mahendran (2008) found that fall-off of plasterboard occurs when the unexposd face of

the board reaches about 1000°C. The temperature at which gypsum boards lose their

restraining capacity depends on the type of board used. However, according to Ranby

(1999) a common temperature of 550°C was proposed. In the numerical study of Kaitila

(2002), the boundary conditions providing lateral restraints at both flanges were

assumed to be valid until 600°C, implying that plasterboard has not fallen off until this

temperature was reached. Thermal properties of gypsum plasterboard are required if

finite element thermal calculations are to be undertaken. Table 2.4 and 2.5 shows the

apparent thermal properties of gypsum board and insulation materials by Alfawakhiri

(2001).

Table 2.4: Apparent Thermal Properties of Firecode C Core Type X Gypsum

Board (12.7mm thick, bulk density 750 kg/m3) (Alfawakhiri, 2001)

Apparent Thermal

Properties

Temperature Range (oC)

<50 50-

80

80-

100

100-

120

120-

140

140-

160

160-

180

180-

200

200-

300

300-

500

500-

700 >700

Conductivity [W /

(moC)]

0.27 0.27 0.27 0.15 0.15 0.15 0.15 0.15 0.17 0.17 0.25 0.45

Heat Capacity [MJ /

(m3o

C)] 0.49 0.7 1.4 2.8 5.6 9.1 7.0 2.8 2.8 1.4 0.49 0.35

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Table 2.5: Apparent Thermal Properties of Insulation Materials (Alfawakhiri,

2001)

Insulation Type

(bulk density in kg/m3)

Apparent

Heat

Capacity

[MJ /

(m3o

C)]

Apparent Thermal Conductivity [W /

(moC)]

In Temperature Range (oC)

<80 80-

200

200-

300

300-

400

400-

500

500-

700 >700

Rock Fibre Batts (33

kg/m3)

0.027 1.0 0.5 0.1 0.1 1.5 2.0 3.0

Glass Fibre Batts (10

kg/m3)

0.009 1.0 0.5 0.1 0.1 1.5 2.0 3.0

Loose Fill Cellulose

(47 kg/m3) 0.115

1.0 0.3 0.3 0.3 1.0 1.0 2.0

(b). Cold-formed Steel

Anderberg (1986) concluded that when modelling material behaviour of steel, a steady

behavioural model can predict the transient test under any given fire process and load.

According to Gerlich et al. (1996) the crystalline structure of carbon steels typically

used in construction changes at temperatures above approximately 650°C. However,

past test results (Alfawakhiri, 2001) indicate that the failure of load bearing LSF

systems is expected to happen before crystalline steel structure changes become a

factor.

Yield Strength

Klippstein (1980b) carried out experimental work on the yield strength of cold-formed

steel framing members as a function of temperature. This data was used to develop the

following equation, which was used by Gerlich et al. (1996).

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11

4

8

3

6

2

4 107.1

109.1

100.4

103.51 TTTT

FyoFyt

(2.1)

where Fyt and Fyo are the yield stresses at temperature T and room temperature,

respectively. Mechanical properties of cold-formed steels decrease rapidly with

increasing temperatures in a fire. Dolamune Kankanamge (2008) has undertaken a study

to investigate the mechanical properties of cold-formed steels at elevated temperatures.

Based on the yield strength results obtained from tensile coupon tests at various

temperatures, a set of equations was developed for low and high strength steels as given

in Equation (2.2) to (2.12).

For low strength steels (G250, G300),

CT o20020 01.10005.020,

, T

f

f

y

Ty (2.2)

CT o800200 022.0

20,

,16.125 T

f

f

y

Ty (2.3)

Equations (2.2) and (2.3) present the proposed equations for reduction factors

( 20,, yTy ff ) of low strength steels, where Tyf , and 20,yf are the 0.2% proof stresses at

elevated and ambient temperatures, respectively, and T is the temperature.

The reduction factors of high strength steels show three main regions: two nonlinear

regions (20oC – 300

oC and 300

oC – 600

oC) and one linear region (600

oC – 800

oC).

Three different equations were therefore developed for these three main regions.

As the first option, Equations (2.4) to (2.6) present the proposed equations for reduction

factors ( 20,, yTy ff ) of high strength steels (G500 and G550).

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For high strength steels (Option 1),

CT o30020

Tx

T

f

f

y

Ty

10

56.4

20,

,

101

201 (2.4)

CT o600300

T

T

f

f

y

Ty

76.7

30095.0

45.1

20,

, (2.5)

CT o800600 35.00004.020,

, T

f

f

y

Ty (2.6)

The equations were developed without considering the results of 0.42 mm G550 steel.

Since 0.42 mm G550 steel is unlikely to be used in load bearing structural members,

this approach is justifiable.

In the second option linear equations for 20oC to 300

oC and 600

oC to 800

oC

temperature ranges and one non-linear curve for 300oC to 600

oC were proposed

(Equations (2.7) to (2.9)).


CT o30020 00358.1000179.020,

, T

f

f

y

Ty (2.7)

CT o600300

T

T

f

f

y

Ty

76.7

30095.0

45.1

20,

, (2.8)

CT o800600 35.00004.020,

, T

f

f

y

Ty (2.9)

As an alternative to Equations (2.4) to (2.9), three simple linear equations were also

developed for the three main regions: 20oC – 300

oC, 300

oC – 600

oC and 600

oC – 800

oC

as given in Equations (2.10) to (2.12).


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CT o30020 00358.1000179.020,

, T

f

f

y

Ty (2.10)

CT o600300 79.10028.020,

, T

f

f

y

Ty (2.11)

CT o800600 35.00004.020,

, T

f

f

y

Ty (2.12)

Elastic Modulus

Similar to yield strength, Klippstein (1980b) obtained experimental data for the

modulus of elasticity for cold-formed steel studs. Gerlich et al.‟s (1996) study fitted a

polynomial to this data which gave an expression as follows

12

4

9

3

7

2

4 104.5

101.6

107.3

100.31 TTTT

EoEt (2.13)

Where Et and Eo are the modulus of elasticity at temperature T and room temperature,

respectively.

New empirical equations were developed for elastic modulus with respect to the

temperature by Dolamune Kankanamge (2008). Deterioration of elastic modulus with

increasing temperature directly influences the performance of the structural member as

it reduces the stiffness. There are two main regions in which reduction factors vary

linearly: 20oC-200

oC and 200

oC–800

oC.

In this study it was found that the influence of steel grade and thickness on the modulus

of elasticity reduction factors is negligible and that there was not any identifiable trend

of reduction of elastic modulus with respect to the steel thickness or grade. Hence

neither steel thickness nor steel grade was included in developing the predictive

equations. Two linear equations were developed for the two identified temperature

Literature Review


regions to predict the elastic modulus reduction factors at elevated temperatures

(Equation (2.14) and (2.15)).

For low and high strength steels,

CT o20020 0167.1000833.020

TE

ET (2.14)

CT o800200 1201.100135.020

TE

ET (2.15)

Coefficient of Thermal Expansion

Steel will expand considerably when exposed to high temperatures. When the steel stud

wall is exposed to fire from one side, thermal bowing will be developed due to the

presence of non-uniform temperatures across the steel section. Hence the knowledge of

the coefficient of thermal expansion is necessary to do the analysis of LSF wall or wall

panels at elevated temperatures.

In the study of load bearing capacity of cold-formed steel joists, Alfawakhiri (2001)

used the following equation for the coefficient of thermal elongation of steel.

e T = ( 0.2x10-8

T2 + 1.2x10

-5 T – 2.408x10

-4 )

Where e T is the coefficient of thermal elongation of steel at temperature T (in ºC) when

heated from 20ºC

Gerlich et al. (1995) in the study of fire resistance of LSF walls used the following

relationship proposed by Lie (1992) for the coefficient of thermal expansion of steel.

αT = (0.004T + 12) x 10-6

, for T < 1000°C

Where αT is the coefficient of thermal expansion at temperature T (°C-1

)

Literature Review


Thermal Conductivity

The temperature rise of a steel member as a result of heat flow is a function of the

thermal conductivity and specific heat of the material. The following equation was used

by Gerlich et al. (1996) to predict thermal conductivity as a function of temperature.

ks = -0.022T + 48, for 0<T<900°C

According to Alfawakhiri (2001), steel framing plays a minor role in the heat transfer

mechanism, hence the accurate determination of thermo-physical properties of steel,

such as specific heat Cs and thermal conductivity Ks is of little importance for the

thermal modelling of LSF walls exposed to fire. Hence an approximate constant value

of 37.5 W/(m°C) was suggested for ks.

Alternative expression for thermal conductivity of steel (ks) is presented in EN 1993-1-2

(NSAI, 2005), where the variation of ks (W/m°C) with the grade of steel is ignored.

ks = 54 – 0.0333 T for T ≤ 800 °C

ks = 27.3 for T ≥ 800 °C

Specific Heat

Lawson and Newman (1990) proposed the variation of Cs (J/kg°C) in relation to

temperature as follows:

Cs = 38.0 * 10-8

T2 + 2.0 * 10

-4 T + 0.47 for 20 °C ≤ T ≤ 725 °C

Literature Review


2.11. Previous Thermal Modelling

The earliest analytical method to predict the time of structural failure for load bearing

LSF walls subjected to standard fire was proposed by Klippstein (1978, 1980). The

following basic assumptions were formulated:

-. The gypsum board does not carry any vertical load.

-. The gypsum board prevents torsional buckling and weak axis (in the plane of the

wall) buckling failure modes of the steel studs.

-. Steel stress-strain relationship at elevated temperatures is linear up to the yield

strength.

-. The total vertical load applied to the wall assembly is always uniformly

distributed among steel studs. The averaged axial load is applied concentrically

to the steel stud section. The studs are hinged at the ends.

-. All studs in the wall assembly experience equal temperature gradients, equal

average temperatures and equal lateral deflections throughout the fire tests. The

temperature gradients and average temperatures are always uniform along the

steel studs.

Klippstein acknowledged the limitations of his method that was heavily dependent on

empirical determination of stud temperatures and lateral deflections.

Gerlich (1995) employed a commercially available computer program, TASEF (Sterner

and Wickstrom 1990), to model heat transfer through LSF walls exposed to fire.

Proprietary (unspecified) thermal properties of gypsum board were used in the

simulations. The numerical predictions showed a good correlation with temperatures

measured in three fire tests. TASEF was reported to yield somewhat non-conservative

stud temperature predictions towards the end of the fire tests. These discrepancies were

attributed to the degradation (opening of joints, cracking, and ablation) of fire exposed

gypsum-board lining which caused an accelerated rise in measured temperature. Gerlich

(1995) used the same basic assumptions (listed earlier) of Klippstein (1978) to analyze

the structural behavior of load-bearing LSF walls. The horizontal mid-height deflection

of the studs was modeled as the sum of two components: stress-free thermal bowing

Literature Review


deflection due to temperature gradient (assumed linear) across the steel stud section, and

secondary deflection due to the average stud load. The mid-height thermal bowing

deflection was treated as the initial eccentricity of the vertical load at stud ends (not at

stud mid-height) (Gerlich et al 1996).

In numerical simulations, stress-free thermal bowing deflections were assumed to

remain constant when temperature difference across the depth of the steel stud

decreased. This was achieved by not allowing the calculated value of the thermal

bowing deflection at any given time step to be less than in the previous step. The total

horizontal deflection at stud mid-height was calculated by adding the thermal bowing

deflection to the secondary deflection (deflection due to the average stud load). The

studs were then analyzed as steel members subjected to axial load and bending moment.

AISI allowable stress formulae, modified to account for the reduction of steel strength

and stiffness at elevated temperatures, were used to estimate the load carrying capacity

of the studs. The failure time was derived as a function of the Load ratio and the

thickness of gypsum board protection. The resulting load-time curves were presented

for a number of generic wall assemblies. These curves are believed to be non-

conservative for high load ratios (Gerlich 1995).

An iteration procedure was used to determine critical temperatures, at which predicted

steel stud capacities were equal to the applied load. These critical temperatures were

then compared with compression flange temperature histories to find failure times.

Gerlich (1995) reported the horizontal deflections calculated from measured

temperatures to agree well with measured mid-height deflections. Failure time

predictions were most accurate when based on measured temperatures. The thermal

model, however, was reported to predict greater than the measured temperature

differences across steel stud sections. Therefore, calculated lateral deflections based on

TASEF temperatures over-estimated the actual mid-height deflections measured in the

fire tests. This effect resulted in slightly conservative failure time predictions but within

80-90% of the test results.

At very high temperatures, some openings of the exposed sheet joints due to deflection

of the framing members and ablation (erosion due to heating) of the exposed linings

allowed hot gases into the cavity. However, these effects of accelerated rise in measured

Literature Review


temperatures towards the end of the tests were not modelled by TASEF. As a result,

differences between predictions and measurements were observed at high temperatures.

Also TASEF does not model mass transfer (moisture movement). Hence the predicted

horizontal deflection using TASEF temperatures exceeded those calculated from

measured temperatures.

A computer program TRACE was developed and used by Alfawakhiri (2001) to

conduct numerical simulations of temperature histories. The thermal properties gained

from literature review were calibrated to produce a good match of numerical and test

results. Hence it was believed that these apparent thermal properties, to some degree,

implicitly account for physical phenomena other than heat transfer, such as mass

transfer, phase change, etc. The presence of the steel frame was neglected in the heat

transfer simulations. The spalling of gypsum boards was modelled by removing it from

the simulation at a user-specified time.

Mathematical and numerical analyses of dehydration of gypsum platerboards exposed

to fire was carried out by Belmiloudi and Meur (2005), and it was found that the

radiative heat transfer between the unexposed surface and the surrounding cannot be

neglected.

Feng et al.(2003a) used the experimental study results of fire tests to validate the

thermal analysis capabilites of ABAQUS. In some of these systems, one or more layers

of gypsum boards on the fire exposed side were removed in numerical studies to

consider possible fall-off of gypsum boards. Also the results of a parametric study using

ABAQUS to examine the thermal performance of steel stud systems with different

numbers of gypsum boards on the exposed and unexposed sides were presented. This

study also assumed a uniform temperature distribution along the stud length. Feng et al.

(2003a) concluded that ABAQUS can be used to simulate the temperature profile in

LSF wall panels under standard fire conditions, including cavity radiation, by adopting

the appropriate thermal boundary conditions and thermal properties, provided there is

no integrity failure of the gypsum boards. It was also found that the temperature profile

of steel stud wall panel was not affected much by the shape of the thin-walled steel

cross-section. The effect of lips on temperature distribution can be ignored, provided

their width is small. It was found that the thermal performance of wall panels was not

significantly affected by the types of interior insulation and the shape of the cold-

Literature Review


formed thin-walled steel cross-section. Temperatures of the steel section of a steel stud

panel system depend primarily on insulation panels on the fire exposed side. However,

it was noticed that, not having interior insulation gave poor fire performance, which is a

contradicting result compared to Kodur and Sultan (2001) and Alfawakhiri (2001).

In the study of Zhao et al. (2005) different computer codes such as ABAQUS, ANSYS,

FLUENT were used to investigate the validity of heat transfer analysis. The results

obtained from these different computer codes showed a good agreement between them

and it was considered that all these computer codes are available for heat transfer

analysis if one of them is validated against tests. It was assumed that conduction is the

main heat transfer mechanism in the steel studs and plasterboards. Convection and

radiation act essentially for heat transfer from fire to plasterboards. As simplification,

radiation effects within the plasterboards were neglected. In numerical models, non-

linearity due to temperature dependency of material properties and boundary conditions

were taken into account. The height and the cross section size of the stud were

considered as parameters affecting the thermal behaviour. However, the mass transfer in

materials such as moisture movement was not simulated. A new set of thermal

properties (specific heat and conductivity) for used plasterboard have been obtained,

after some numerical investigations, which led to a good estimation of stud heating

compared to test results. However, these proposed thermal properties were not

presented in this report.

Sultan (1994) found that the temperature distribution in the gypsum panels, even

relatively close to the steel studs, were substantially one-dimensional through the

thickness of the panels. This finding and the experimentally validated one-dimensional

thermal response models of Sultan (1996) and Cooper (1997) indicate that a

compartment fire model whose model equations include a one-dimensional heat transfer

analysis for gypsum-panel/steel-stud wall system thermal response can lead to an

accurate overall accounting of energy conservation, and can yield accurate wall system

thermal response simulations even up to the time of failure.

In spite of the above finding, heat transfer through gypsum-panel/steel-stud wall

systems is not a totally one-dimensional phenomenon. In particular, near the “sparse”

regions of the stud/gypsum-panel joints, the heat transfer problem are strongly two-

dimensional, i.e., an accurate determination of the steel stud thermal response will

Literature Review


require a two-dimensional time-dependent analysis (with two materials, steel and

gypsum) of these regions. Furthermore, for load-bearing type wall systems and in terms

of the critical evaluation of wall system fire resistance, it is the spatially varying loss of

strength of the steel studs due to spatially varying elevated temperatures that would lead

to possible wall system structural failure, and that would have to be simulated.

Nevertheless (Figure 2.12), the use of a one-dimensional analysis in the overall fire

model equation can lead to accurate simulations of room fire environments. One

dimensional analysis could also be used to create an accurate simulation of wall system

thermal response if that section of the wall is far away from steel studs and tracks.

Figure 2.12: Sketch of the idealized geometry of the gypsum-panel/steel-stud wall

system

The thermal/structural fire model, SAFIR, developed and currently being advanced at

the University of Liege, has been identified and used as a reliable thermal/structural

computational model. Preliminary applications of SAFIR at the Centre of Advanced

Technology for Large Structural Systems (ATLSS) of Lehigh University and the

University of Maryland indicate that the above approach to coupling a compartment fire

model and a thermal/structural fire model will be successful.

Literature Review


This brief review of thermal modelling shows that it is possible to obtain reasonably

accurate results from thermal modelling of LSF wall systems using the many different

numerical tools available to fire researchers despite the complex LSF system and

behaviour and associated simplifications in modelling.

2.12. Heat Transfer Simulation

The presence of steel frame was neglected in heat transfer simulations, because due to

the light weight of thin gauge members, they play a minor role in the heat transfer

mechanism. Alfawakhiri (2001) used a large number of numerical trial runs to arrive at

a single set of thermal properties that would produce a reasonable agreement of

simulated and measured temperature histories. In other words, the material properties

were essentially calibrated to produce a good match of numerical and test results. The

calibration process was to start numerical simulations with a set of thermal properties,

reported in the literature, and compare the output histories with measured temperatures.

Then after changing a thermal property in a narrow temperature interval, one at a time, a

trial numerical run would be conducted and the output results checked for a better

agreement with experimental data. This procedure was repeated many times first to

calibrate the properties of gypsum board to match measured temperature histories of

tests (non-insulated wall test of the longest duration with all gypsum board staying in

place until the end of the test). Secondly, these gypsum board properties were verified

in the simulation of tests (non-insulated wall tests of shorter duration with observed fall-

off of the face layer of the gypsum board). The third step was to apply the calibrated

gypsum board properties in the simulations of tests (insulated walls) and calibrate the

properties of insulation materials (rock fibre batts, glass fibre, and loose fill cellulose).

The apparent material properties, thermal conductivity and heat capacity at temperatures

up to 1000°C, were found to have a great deal of influence on the shape of simulated

time-temperature curves. It should be mentioned, however, that these apparent thermal

properties, to some degree, implicitly account for physical phenomena other than heat

transfer, such as mass transfer, phase change, etc. This happens because the temperature

rise in LSF walls exposed to fire is affected by processes not described by heat transfer,

such as migration of moisture vapors, penetration of cool ambient air or hot furnace

gases into the cavity, etc.

Literature Review


Another parameter that has a major effect on simulated temperature histories is the fall-

off time of gypsum board layers. TRACE models the spalling of gypsum board by

removing it from the simulation at a user-specified time. The fall-off times is the

beginning of layer spalling based on visual test observations. In the retrospective

simulations, these times were slightly adjusted in order to represent a time when a

significant portion of the layer had fallen off. In both simulated and measured

temperature histories, the fall-off of gypsum board layers is usually manifested in

respective time-temperature curves by sudden shifts in temperature closely approaching

the furnace temperature.

2.13. Thermal Performance

The conventional method to increase the fire rating of LSF wall system is simply by

adding more plasterboards or place insulation materials inside the cavity. Kolarkar and

Mahendran (2009) conducted some full fire scale tests to understand more about the

current design practice and make further development. In their tests (Table 2.3) adding

more plasterboard doubled the fire rating from 53 min to 111 min. Table 2.6 shows the

summary of FRM fire resistance tests on load bearing LSF walls.

The Fire Risk Management Program (FRM) of the Institute for Research in

Construction (IRC) of the National Research Council of Canada (NRC) conducted some

fire tests of conventional LSF wall systems. The wall assemblies tested (designated as

WI, W2 and W3) were 3048 mm high, 3658 mm long and 157 mm deep. Each assembly

consisted of a single row of galvanized cold-formed steel studs, protected with two

layers of fire resistant gypsum board on each side. The failure times from both tests

(Kolarkar, 2009 and FRM fire test) are roughly the same, around 55 minutes.

Literature Review


Table 2.6: Summary of FRM Fire Resistance Tests on Load Bearing LSF Walls

Specimen

number

Stud

spacing

(m)

Insulation

type

(fibre)

Resilient

channels

on

exposed

side

Load

including

self

weight

(kN/m)

Fall-off time of

gypsum board

on exposed side

(min)

Structural

failure

time

(min)

Temperature rise on

unexposed side,

under pads, at failure

time (oC)

Face

layer

Base

layer

Maximum Average

W1 406 Glass Yes 21.5 50

in

place 55 52 36

W2 610 Rock Yes 14.3 57 67 73 50 42

W3 406 Cellulose Yes 21.5 57

in

place 70 42 37

Adding insulation material inside the cavity did not increase the fire rating considerably

thus Kolarkar and Mahendran (2009) developed a new composite LSF system. Instead

of having insulation material inside the cavity, they placed it between plasterboards

(external insulation). This method increased the fire rating of the system dramatically.

Alfawakhiri (2001) also mentioned that insulation placed in the cavity of load bearing

LSF walls reduces their fire resistance because it reduced the ability of fire exposed

gypsum board to remain in place and because it causes non-uniform heating of the load

bearing steel studs. From these full scale fire tests Kolarkar and Mahendran (2009)

generate the standard Time-Temperature Curves for the new LSF composite wall

system. But further analytical model is needed to develop idealized Time-Temperature

profiles for standard ISO curve or even realistic fire curve for all LSF walls. In Figures

2.13 and 2.14, the sudden spike in steel stud temperature was caused by the spalling of

gypsum boards.

Literature Review


Figure 2.13: Typical Temperature History of a Steel Stud within an LSF Wall

(Alfawakhiri, 2001)

Figure 2.14: Time-Temperature Curves for Glass Fibre Composite Panel

(Kolarkar, 2010)

2.14. SCI Publication (SCI, 1993)

SCI (1993) presents guidance for the fire resistance of protected sections in floors or

walls acting as compartment boundaries, i.e. planar protection. In this case, heat is

Time-Temperature Graphs (NLB:Sp7-2x2-Composite Panel-GF)

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190 200 210 220

Time (min)

Tem

pera

ture

(oC

)

AS 1530.4 Furnace FS Pb1,Ins Ins,Pb2

Pb2 Cav Pb3 Cav Pb3,Ins Ins,Pb4 AS

Literature Review


applied from one side only and the floors or walls must satisfy the necessary insulation

criterion. The thickness of fire protection for conventional beams and column is

determined by using the method given in BS 5950: Part 8 for extending the existing data

for hot-rolled sections to cover the use of cold-formed steel sections. Design tables are

presented for typical materials and section sizes. Table 4 of this publication defines the

fire resistance time in respect of different parameters such as the number of

plasterboards, protection thickness, type of plasterboard and insulation. This table is

reproduced as Table 2.7 in this thesis.

Table 2.7: Fire Resistance of Typical Floors, Walls and Partitions Comprising

Cold-formed Steel Sections and Planar Board Protection (heated from one side

only) (SCI, 1993)

Form of

construction

Number

of layers

of board

Protection

thickness

(mm)

Fire Resistance (hours) Notes

Plasterboard Fire

resistant

board†

Floors with

ceiling

protection

1 12.5 - 0.5 -

2

2

12.5

15

0.5

-

1.0

1.5

+ 60 mm glass

wool mat**

-

Non-load

bearing

walls

(partitions)

(number of

layers per

face)

1

1

1

12.5

12.5

15

0.5

0.5

0.5

0.5

1.0

1.0

-

+ 25 mm glass

wool mat*

-

2

2

12.5

12.5

1.0

1.0

1.5

2.0

-

Boxed section

Literature Review


2

15

1.5

2.0

depth > 60 mm

-

Load bearing

walls

1 12.5 - 0.5 -

2

2

12.5

15

0.5

-

1.0

1.5

-

-

† „Fireline‟ or „Firecheck‟ board or similar

* Glass wool mat is required for insulation purposes for more than 30 minutes fire

resistance

** For floors, the glass wool mat is only necessary for fire resistant suspended ceilings

The strength of cold-formed steel that may be used in the calculations at the fire limit

state is presented in Table 3 of SCI,1993. These strength reduction factors are expressed

as a ratio of the normal (room temperature) strength, and are based on the 95%

confidence limit. This table is reproduced in this report as Table 2.8. These data are

used in establishing the limiting temperatures of cold-formed sections used as structural

members. Table 2.9 reproduces these limiting temperatures in this report.

Table 2.8: Strength Reduction Factors for Cold-formed Steel at Elevated

Temperatures (SCI, 1993)

Temperature (°C) 200 250 300 350 400 450 500 550 600

0.5% strain 0.95 0.89 0.83 0.76 0.68 0.58 0.47 0.37 0.27

1.5% strain 1.00 0.99 0.95 0.88 0.82 0.69 0.56 0.45 0.35

Literature Review


Table 2.9: Limiting Temperature (oC) of Beams and Columns using Cold-formed

Steel Sections (SCI, 1993)

Member Type

Load ratio at fire limit state

0.7 0.6 0.5 0.4 0.3

Beams supporting concrete slabs 530 555 600 640 670

Beams supporting timber floors 450 485 530 575 625

Columns in walls 445 480 520 560 605

Slender columns 400 450 490 540 590

Other elements; Studs and ties 400 450 490 540 590

2.15. Literature Review Findings

Following are the main findings from this literature review.

Panel Systems

Recently Kolarkar and Mahendran (2008) developed a new composite panel

system for LSF walls and floors, where the insulation was placed outside the

steel frame and it was found that the fire resistance has improved considerably.

However, Kolarkar and Mahendran‟s (2008) study was limited to an

experimental study with a load ratio of 0.2. Gunalan (2009) has extended this

work by conducting three fire tests for a load ratio of 0.4 and commenced

structural modelling of tested walls. Hence further numerical and theoretical

analyses are needed to fully understand the improvements offered by the new

system to its thermal performance.

Thermal Modelling, Thermal Performance and Material Properties

The thermo-physical and mechanical properties of most materials change

substantially within the temperature range associated with building fires. In the

Literature Review


field of fire science, applied materials research faces numerous difficulties. At

elevated temperatures, many building materials undergo physicochemical

changes. Most of the properties are temperature dependent and sensitive to

testing method parameters such as heating rate, strain rate, temperature gradient,

etc. There has been a tendency to use “notional” (“typical”, “proprietary”,

“empirical”, etc.) values for material properties in numerical computations, in

other words, values that ensure agreement between experimental and analytical

results.

Alfawakhiri (2001) used the thermal properties gained from literature review

and calibrated them to produce a good match of numerical and test results. It

was believed that these apparent thermal properties, to some degree, implicitly

account for physical phenomena other than heat transfer, such as mass transfer,

phase change, etc. The presence of the steel frame was neglected in the heat

transfer simulations, because, due to the light weight of thin gauge members,

they play a minor role in the heat transfer mechanism.

Another parameter that has a major effect on simulated temperature histories is

the fall-off time of gypsum board layers. The spalling of gypsum boards was

modelled by removing it from the simulation at a user-specified time. The fall-

off time is the beginning of layer spalling based on visual test observations. In

the retrospective simulations, these times were slightly adjusted in order to

represent a time when a significant portion of the layer had fallen off. In both

simulated and measured temperature histories, the fall-off of gypsum board

layers is usually manifested in respective time-temperature curves by sudden

shifts in temperature closely approaching the furnace temperature.

In the study of Zhao et al. (2005) different computer codes such as ABAQUS,

ANSYS and FLUENT were used to investigate the validity of heat transfer

analysis. The results obtained from these different computer codes showed a

good agreement between them and it was considered that all these computer

codes are available for heat transfer analysis if one of them is validated against

tests. It was assumed that conduction is the main heat transfer mechanism in the

steel studs and plasterboards of LSF walls. Convection and radiation act

essentially for heat transfer from fire to plasterboards. As simplification,

radiation effects within the plasterboards were neglected. In numerical models,

Literature Review


non-linearity due to temperature dependency of material properties and

boundary conditions were taken into account. The height and the cross section

size of the stud were considered as parameters affecting the thermal behaviour.

However, the mass transfer in materials such as moisture movement was not

simulated.

Physical experiments have shown that as the porosity of glass fibre insulating

material increases, the amount of glass fibre decreases and the effective heat

transfer coefficient of the material also decreases. This is in accordance with the

fact that the coefficient of thermal conductivity of glass is quite large compared

to that of air.

Mathematical and numerical analyses of dehydration of gypsum plasterboards

exposed to fire was carried out by Belmiloudi and Meur (2005), and it was

found that the radiative heat transfer between the unexposed surface and the

surrounding cannot be neglected.

In numerical simulations, stress-free thermal bowing deflections were assumed

to remain constant when temperature difference across the depth of the steel stud

decreased. The total horizontal deflection at stud mid-height was calculated by

adding the thermal bowing deflection to the deflection due to the average stud

load.

An iteration procedure was used to determine the applied critical temperatures,

at which predicted steel stud capacities were equal to the applied load. These

critical temperatures were then compared with compression flange temperature

histories to find failure times.

The temperature distribution in the flat elements such as walls, floors, and

ceilings is essentially one-dimensional, with a gradient only across the thickness.

Opening of joints, cracking, and ablation of fire exposed gypsum-board lining

caused an accelerated rise in measured temperatures.

The thermal performance of a steel cassette wall system is greatly affected by its

layout. In order to achieve good fire resistance, the narrow webs of the cassette

sections should be put on the fire exposed surface. This will help reduce the

amount of heat attracted by the cassette system.

Kolarkar (2009) and Gunalan (2009) have produced extensive thermal

performance data of a range of LSF wall systems in terms of time-temperature

Literature Review


profiles under standard fire conditions. However, there is a need to develop a

validated numerical model that can simulate the observed time-temperature

profiles. Once a validated numerical model is available a combination of

experimental and numerical results can be used to develop time-temperature

profiles for various LSF wall systems in particular for the new LSF wall system

with a composite panel. There is also a need to investigate the effect of various

LSF wall components on their thermal performance under both standard and

more realistic fire conditions.

Structural Performance

Basic assumptions used in the prediction of structural failure of LSF walls

(Klippstein, 1978, Alfawakhiri, 2001) were as follows :

-. The gypsum board does not carry any vertical load.

-. The gypsum board prevents torsional buckling and weak axis (in the

plane of the wall) buckling failure modes of steel studs.

-. Steel stress-strain relationship at elevated temperatures is linear up to the

yield strength.

-. The total vertical load applied to the wall assembly is always uniformly

distributed among steel studs. The averaged axial load is applied

concentrically to the steel stud section. The studs are hinged at the ends.

-. All studs in the wall assembly experience equal temperature gradients,

equal average temperatures and equal lateral deflections throughout the

fire tests. The temperature gradients and average temperatures are always

uniform along the steel studs.

AISI allowable stress formulae, modified to account for the reduction of steel

strength and stiffness at elevated temperatures, were used to estimate the load

carrying capacity of the studs. The failure time was derived as a function of the

load ratio and the thickness of gypsum board protection.

Literature Review


Screw spacing from the gypsum board edge, cavity insulation, and the number

of gypsum boards had a significant effect on the fire performance of wall

assemblies.

The non-uniformity of load distribution among studs, the end conditions of the

studs, the concentricity of load application and the role of the gypsum board in

the structural behavior of LSF walls in standard fire tests are the many issues

that require further investigation.

Much higher fire resistance ratings are likely to be achieved in fire tests on

similar non-load bearing LSF assemblies with lower loading ratio since loading

plays a significant role in fire resistance tests.

Experimental Study of Thermal Properties


Chapter 3


3.1. General

In this study three materials used commonly in the construction of LSF wall systems are

considered. They are gypsum plasterboard, glass fibre and rock fibre insulations.

Thermal properties of gypsum board and insulation materials are very important

parameters needed in the numerical modelling of LSF wall systems. These thermal

properties are density, thermal conductivity, and specific heat.

Since these properties are very sensitive to the changes in temperature, it is very

important to measure them as a function of temperature. Thermo Gravimetric Analysis

(TGA) method was used to determine the effects of dehydration/calcinations and

decomposition of these thermal properties. A Differential Scanning Calorimeter (DSC)

machine was used to determine the specific heat and mass loss as a function of

temperature and the apparatus used was SETARAM TGA DSC.

This chapter presents the details of an experimental study undertaken to determine the

specific heat and mass loss of plasterboard, glass fibre and rock fibre insulations. These

measurements were then used to compile an experimental thermal property database

required to validate the numerical models that could be used to predict their thermal

performance under standard fire tests (AS 1530.4).

3.2. Test Specimens

The gypsum plasterboard tested was BORAL Firestop. It is a widely used brand in

Australia for constructing LSF wall panels. Its dimensions are 2400 mm x 1800 mm

with 16 mm thickness. Regular or standard board is not required to have any fire

resistant rating, so it usually has a low density gypsum core with no reinforcing except

the external paper. The glass fibre tested was BORAL Insulation. Like Boral Firestop, it

is also a widely used glass fibre insulation in Australia. The rock fibre tested was



Tombo Brand M.G. Mighty Rock Fibre. It is a flexible rock fibre blanket for insulation

of roofing, wall, instrument piping, elbows and ducts, etc.

According to ASTM E 1269 (ASTM, 2005), specimen materials need to be ground to a

powder form and have at least 6 to 12 mg in mass to conduct the required DSC test.

Table 3.1 shows the initial mass of each specimen used in this study. A total of eight

specimens was used as seen in Table 3.1. Powdered or granular specimens were mixed

prior to sampling and should be sampled by removing portions from various parts of the

container. These portions, in turn, were combined and mixed to ensure a representative

specimen for the determination.

Table 3.1: Initial Mass of Materials Used in the DSC Test

Gypsum Plasterboard Glass Fibre Rock Fibre

Specimen Mass (mg) Specimen Mass (mg) Specimen Mass (mg)

1 11.30 1 11.30 1 11.30

2 11.09

3 11.41 2 11.09 2 11.09

4 11.24

To satisfy ASTM E 1269 (ASTM, 2005) requirements of powder samples, the QUT

grinding machine shown in Figure 3.1 was used. The machine consists of a steel tube

with three steel balls inside it. The solid material specimen was put inside the tube and

vibrated for about 5 to10 minutes in order to grind the sample to a powder form. Figures

3.2 to 3.4 show the test specimens after grinding.

Figure 3.1: QUT Grinding Machine



Figure 3.2: Plasterboard in Powder Form

(a) Before Grinding (b) After Grinding

Figure 3.3: Glass Fibre Insulation

(a) Before Grinding (b) After Grinding

Figure 3.4: Rock Fibre Insulation

Thermo Gravimetric Analysis (TGA) needs a reference material to calculate the specific

heat of the sample. In this experiment Al2O3 was used as the reference material. Table

3.2 shows the initial mass of Al2O3 used in the tests. ASTM E 1269 (ASTM, 2005)



provides the standard Al2O3 specific heat (Cp) values required for the calculation of the

sample specific heat.

Figure 3.5: Al2O3 Powder

Table 3.2: Al2O3 Used in the DSC Test

DSC TEST Al2O3 (mg)

Gypsum Board Specimen 1 13.92




Glass Fibre Specimen 1 11.71

Glass Fibre Specimen 2 11.71

Rock Fibre Specimen 1 11.71

Rock Fibre Specimen 2 11.71

3.3. Test Set-up and Procedure

Since a host of measurement methods has been used over the years to quantify thermal

properties, it is very difficult to understand if differences in the reported measurements

are due to material properties or the varying methods employed. In this study ASTM E

1269 (ASTM, 2005) guidelines were used to conduct the DSC tests. The apparatus used

to conduct the test was SETARAM TGA DSC as shown in Figure 3.6.

Figure 3.6: SETARAM TGA DSC



Four DSC tests were conducted on gypsum plasterboard, two for glass fibre insulation,

and two for rock fibre insulation. Each of the powdered form specimens was mixed and

inserted into 100 µm aluminium crucibles. Other researchers (Wakili et al, 2006) used a

lid on the crucible to investigate the evaporation effect of specimens. In this study, a lid

was not used because it would stick with the aluminium crucible when the sample was

heated, making the cleaning process very difficult. The aluminium crucible with the

sample inside it was put at the front of SETARAM TGA DSC and the blank aluminium

crucible was put at the back as a correction factor (see Figure 3.7).

Figure 3.7: Aluminium Crucible Configurations

SETARAM TGA DSC need water and gas to run the experiment (see Figure 3.8).

Typical purge gases are air, helium and nitrogen. They are very efficient for heat

transfer and removal of volatiles. In this study nitrogen was used to accommodate the

heat transfer, removal of volatiles and gas generation incurred from dehydration.

Figure 3.8: Water and Nitrogen Control Knobs



Measurements were performed under air flow starting at 20oC up to 550

oC at a heating

rate of 20oC/min under a constant nitrogen gas flow with 100 kPa pressure and a flow

rate between 1.5L – 2L per hour. It was possible to perform a measurement up to

1000oC but titanium crucibles are needed instead of aluminium crucibles. In addition,

the specific heat of blank crucible as a correction material was measured under the same

operating conditions in order to obtain a correction factor. This procedure was used for

plasterboard, glass fibre and rock fibre tests. The same procedure was used for the

reference material, Al2O3. Since heat flow calibration was carried out by using a

standard reference material (Al2O3), continuous specific heat with reference method was

used to calculate the specific heat of the sample (Equation 3.2).

3.4. Analysis of Experimental Results

3.4.1. Typical Experimental Results of Each Specimen

SETARAM TGA DSC provides heat flow output in each test as a function of

temperature and time. These outputs can then be used to obtain the specific heat (Cp).

Figures 3.9 to 3.11 show the typical DSC test results for gypsum plasterboard, glass

fibre and rock fibre insulations.

Figure 3.9: Typical DSC Results of Heat Flow versus Time for Gypsum

Plasterboard

0

150

300

450

600

-150

-100

-50

0

50

100

150

0 400 800 1200 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)

Blank HF Reference HF Plasterboard HF DSC Temperature



Figure 3.10: Typical DSC Results of Heat Flow versus Time for Rock Fibre

Figure 3.11: Typical DSC Results of Heat Flow versus Time for Glass Fibre

0

150

300

450

600

0

50

100

150

0 200 400 600 800 1000 1200 1400 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)

Rock Fibre HF Blank HF Reference HF DSC Temperature

0

150

300

450

600

0

50

100

150

0 200 400 600 800 1000 1200 1400 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)

Glass Fibre HF Blank HF Reference HF DSC Temperature



3.4.2. Calculation Methods

There are several methods to calculate the specific heat from the heat flow data of the

DSC experiments.

1. Continuous Cp without reference

In this case, the precise determination of Cp requires two tests with the same

experimental conditions:

-. The first test is conducted with two empty vessels without the sample (blank).

-. The second test is conducted with the vessels and the sample.

The difference between the two signals is proportional to the specific heat of the

sample. This magnitude is converted directly into thermal power by the calibration

curve of the DSC. This method supplies the determination directly for each temperature

and does not use a reference sample such as Al2O3. Figure 3.12 shows typical DSC

results experiment without reference material

Figure 3.12: Typical Continuous Cp DSC Results without Reference Material

The formula is given next:

dt

dTMassTySensitivit

HFHFTC

sample

blanksample

p

).(

)(

(3.1)



where:

Cp(T) = Sample Specific heat per unit mass at a given temperature

HFsample = Heat flow from the sample at a given temperature

HFblank = Heat flow from the blank crucible at a given temperature

Sensitivity(T) = Apparatus sensitivity at a given temperature

Mass sample = Mass of the sample

2. Continuous Cp (mass) with reference

In this case, the precise determination of Cp requires three tests with the same



-. The second test is conducted with vessels and the reference sample in one of them.

-. The third test is conducted with vessels and the sample in one of them.

The reference sample is substance (Al2O3) whose Cp value is known (see Figure 3.13).

Figure 3.13: Specific Heat of Al2O3 from ASTM E 1269(ASTM, 2005)

0

200

400

600

800

1000

1200

1400

-200 -100 0 100 200 300 400 500 600 700 800

Spe

cifi

c H

eat

(J/

kg.o

C)

Temperature (oC)



Figure 3.14: Typical Continuous Cp (mass) with the DSC Result of Reference

Material


)(..)( TCMass

Mass

HFHF

HFHFTC

refp

sample

ref

blankref

blanksample

p

(3.2)

where:

Cp(T) = Sample Specific heat per unit mass at a given temperature



HFref = Heat flow from the reference crucible (Al2O3 for this study) at a given

temperature


Mass ref = Mass of the reference material (Al2O3 for this study)

Cpref(T) = Reference Specific heat per unit mass at a given temperature (Al2O3 Cp

from ASTM E 1269 for this study – Figure 3.13)

Figure 3.14 shows typical DSC results experiment with reference material. In this case

Al2O3 act as a reference material which specific heat per unit mass in known.



3. Continuous Cp (volume) with reference

In this case, the precise determination of Cpv requires three tests with the same



-. The second test is conducted with vessels and the reference sample in one of them.

-. The third test is conducted with vessels and the sample in one of them.

The reference sample is a substance whose Cpv equation is known.

Figure 3.15: Typical Continuous Cpv (volume) with the DSC Result of Reference

Material


)(.)( TCHFHF

HFHFTC

refpv

blankref

blanksample

pv

(3.3)

Where:

Cpv(T) = Sample Specific heat per unit volume at given temperature



HFref = Heat flow from the reference crucible at a given temperature




Mass ref = Mass of the reference material

Cpvref(T) = Reference Specific heat per unit volume at a given temperature

Figure 3.15 shows typical DSC results experiment with reference material. In this case

Al2O3 act as a reference material which specific heat per unit volume in known.

3.4.3. Results for Plasterboard

The specific heat versus temperature curves were determined from the heat flow

temperature versus time curves using Method 2 (continuous Cp (mass) with reference).

The Heat Flow curves from Specimen 1 to Specimen 4 in Figure 3.16 indicate two

dehydration/calcination steps and decomposition of CaSO4 starting from 130oC to

200oC, which agrees well with the weight loss curve in Figure 3.17. During this

temperature period, two peaks were observed; one at 140oC and the other at about

170oC (see Figure 3.18). The first peak specific heat value varies between 17500 to

22000 J/(kg.oC) while the second peak specific heat value varies between 13000 to

17000 J/(kg.oC) (see Figure 3.18). At about 400°C, a third, exothermic reaction occurs,

in which the molecular structure of the soluble crystal restructures itself into a lower

insoluble energy state (see Figure 3.18). This observation is simliar to Manzello et al.’s

(2008) findings. Figure 3.17 shows the measured mass loss of gypsum plasterboard

(relative density) as a function of temperature. It shows the large decrease in the density

start at approximately 125°C where the first dehydration reaction occurs. In this time

the mass of the plasterboard is reduced by approximately 10% in all cases, however

there is a slight difference in mass loss in each case. This difference in weight loss is

due to the difference in heat flow from each specimen.

It was assumed that the water disappears instantaneously from the board, neglecting the

heat transfer associated with the hot vapour migrating in the gypsum plasterboard from

the fire to the room side. This assumption is due to the small thickness of the gypsum

boards, and will not be applicable to a lightweight wall construction consisting of two

gypsum boards separated by an insulation material. It is most probable that in such a



case the hot vapour emerging from the gypsum board on the fire side will easily migrate

through the insulation layer and condense on the inner surface of the room side gypsum

layer causing a remarkable heat transport from the fire side to the room side of the wall

(Ghazi et al., 2006).

(a) Specimen 1

(b) Specimen 2

Figure 3.16: Heat Flow versus Time for Plasterboards

0

150

300

450

600

-150

-100

-50

0

50

100

150

0 400 800 1200 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)

Blank Reference Plasterboard 1 Temperature

0

150

300

450

600

-150

-100

-50

0

50

100

150

0 400 800 1200 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)




(c) Specimen 3

(d) Specimen 4

Figure 3.16: Heat Flow versus Time for Plasterboards

0

150

300

450

600

-150

-100

-50

0

50

100

150

0 400 800 1200 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)


0

150

300

450

600

-150

-100

-50

0

50

100

150

0 400 800 1200 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)




Figure 3.17: Mass Loss in Plasterboards

Figure 3.18: Specific Heat of Plasterboards

0

1

2

3

4

5

6

7

8

9

10

11

12

0 100 200 300 400 500

Mas

s (m

g)

Temperature (oC)

Plasterboard 1 Plasterboard 2 Plasterboard 3 Plasterboard 4

-2500

2500

7500

12500

17500

22500

0 100 200 300 400 500

Spe

cifi

c H

eat

(J/

kg.o

C)

Temperature (oC)

Plasterboard 1 Plasterboard 2 Plasterboard 3 Plasterboard 4

First Peak at 140oC

Second Peak at 170oC

125



3.4.4. Results for Rock Fibre Insulation

Figure 3.19 shows the rock fibre sample after the DSC test while Figure 3.20 presents

the heat flow versus time curves. The heat flow curves from Specimen 1 and 2 indicate

that the heat flow decreases linearly and no sudden peak occurred unlike plasterboard

heat flow. The heat flow starts to decrease from 75oC to the end of the test at 550

oC (see

Figure 3.20). Rock fibre has a high melting point at more than 1000oC and since the test

was stopped at 550oC any peak in specific heat and mass loss was not observed. The

rock fibre itself was completely burnt and a change in colour was observed. From the

two DSC tests, the specific heat of the rock fibre was found to be in the range of 800

J/kg.oC to 1100 J/kg.

oC as shown in Figure 3.22.

Figure 3.19: Rock Fibre After DSC Test

Figure 3.20: Heat Flow versus Time for Rock Fibre Insulation

0

150

300

450

600

0

50

100

150

0 200 400 600 800 1000 1200 1400 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)

Rock Fibre 2 Blank Reference Rock Fibre 1 Temperature



Figure 3.21: Mass Loss in Rock Fibre Insulation

Figure 3.22: Specific Heat of Rock Fibre Insulation

0

2

4

6

8

10

12

14

16

18

20

0 100 200 300 400 500 600

Mas

s (m

g)

Temperature (oC)

Rock Fibre 1 Rock Fibre 2

0

1000

2000

3000

4000

5000

50 100 150 200 250

Spe

cifi

c H

eat

(J/

kg.o

C)

Temperature (oC)

Rock Fibre 1 Rock Fibre 2



3.4.5. Results for Glass Fibre Insulation

Figure 3.23 shows the glass fibre sample after the DSC test while Figure 3.24 presents

the heat flow and temperature versus time curves. The heat flow curves from Specimen

1and 2 indicate that the heat flow decreases linearly and no sudden peak occurred unlike

the plasterboard heat flow. The heat flow starts to decrease from 100oC to the end of the

test at 550oC (see Figure 3.24). Glass fibre melts at about 700

oC and cannot withstand

direct fire exposure (Sultan and Lougheed 1994). Since the test was stopped at 550oC

any peak in specific heat and mass loss was not observed. The glass fibre itself was

completely burnt and a change in colour occurred. From the two DSC tests, the specific

heat of the glass fibre was found to be between 250 J/kg.oC and 1000 J/kg.

oC as shown

in Figure 3.26.

Figure 3.23: Glass Fibre after DSC Test

Figure 3.24: Heat Flow versus Time for Glass Fibre Insulation

0

150

300

450

600

0

50

100

150

0 200 400 600 800 1000 1200 1400 1600

Tem

pe

ratu

re (

oC

)

He

at F

low

(µ

V)

Time (s)

Glass Fibre 2 Blank Reference Glass Fibre 1 Temperature



Figure 3.25: Mass Loss in Glass Fibre Insulation

Figure 3.26: Specific Heat of Glass Fibre Insulation

0

2

4

6

8

10

12

14

16

18

20

0 100 200 300 400 500 600

Mas

s (m

g)

Temperature (oC)

Glass Fibre 1 Glass Fibre 2

0

1000

2000

3000

4000

5000

50 100 150 200 250 300 350 400 450 500

Spe

cifi

c H

eat

(J/

kg.o

C)

Temperature (oC)

Glass Fibre 1 Glass Fibre 2



3.5. Idealised Thermal Properties to be used in Numerical Models

The results produced from the DSC tests may not be applied directly for modelling

purposes since the LSF wall panels in full scale fire tests used materials with paper and

reinforcement in it, not powder form material. Therefore a large number of numerical

analysis was conducted to arrive at a single set of idealised apparent thermal properties

that would produce a reasonable agreement of simulated and measured temperature

histories. In other words, the thermal properties shown in Figures 3.28 to 3.34 were

essentially calibrated to produce a good match of numerical and test results of LSF wall

panels under fire conditions.

The calibration process was to conduct numerical simulations and compare their output

results with corresponding measured temperature profiles from the fire tests. After

changing a thermal property in a narrow temperature interval, one at a time, a trial

numerical run was conducted and the output results checked for a good agreement with

experimental data. This procedure was repeated many times by changing the thermal

properties of gypsum board until a good agreement was obtained with the measured

temperature histories of small scale tests of Specimens 1 to 5 in Kolarkar (2010).

Secondly, these gypsum plasterboard properties were verified in the simulation of load

bearing wall Tests 1 to 3. The procedure to extract the idealised thermal properties of

gypsum plasterboard are summarised in Chart 3.1. Figure 3.27 presents the specific heat

values suggested by other researchers while Figure 3.28 and Table 3.3 show the

idealised apparent specific heat values of plasterboard. For the specific heat, used as an

input for the numerical analysis in this thesis, the energy consumption due to the double

step dehydration was added. The fist peak was at 140oC with 17500 J/kg.K specific heat

and the second peak was at 170oC with 13000 J/kg.K specific heat (see Figure 3.28). As

shown in Figure 3.28, decomposition process starts at 670oC with 3000 J/kg.K peak.

Overall, the specific heat properties correspond well with the specific heat values

suggested by SAFIR (2004). Figures 3.29 and 3.30 and Table 3.4 provide the idealised

thermal conductivity and mass loss of gypsum plasterboard. The validation of finite

element models using the idealised thermal properties of plasterboard will be discussed

in Chapters 4 and 5.

After the gypsum plasterboard thermal properties were finalised, the next step was to

apply the idealised gypsum plasterboard properties in the simulation of small scale Test



Specimens 6 to 11 (glass and rock fibre insulation) and calibrate the thermal properties

of insulation materials. Once a good agreement between the experimental tests and

numerical results was achieved, these thermal properties were used in the simulation of

the load bearing wall Tests 4 to 7 with cavity and external insulations. The procedure to

extract idealised thermal properties of glass and rock fibre insulations are summarised in

Chart 3.2. Figures 3.31 to 3.34 show the idealised thermal properties for glass and rock

fibre insulations. They indicate that the specific heat of glass fibre is 900 J/kg.oC while

it is 840 J/kg.oC for rock fibre.

Satisfactory Unsatisfactory

agreement agreement


agreement agreement

Chart 3.1: Process to Determine the Idealised Thermal Properties of Gypsum

Plasterboard

Measure specific heat and mass loss of gypsum plasterboard

Use specific heat and mass loss from DSC test results and thermal conductivity from

previous research work to perform finite element modelling using SAFIR

Compare small scale gypsum plasterboard

experimental results from Kolarkar (2010)

with numerical results (Specimens 1 to 5)

Compare load bearing wall experimental

results from Kolarkar (2010) with numerical

results (Tests 1 to 3)

Idealised thermal properties of gypsum plasterboard are established and can be used

in further finite element modelling with gypsum plasterboard

Revise thermal properties

Re-run SAFIR

Perform DSC tests of gypsum plasterboard




agreement agreement


agreement agreement

Chart 3.2: Process to Determine the Idealised Thermal Properties of Rock and

Glass Fibre Insulations

For all the materials, thermal conductivity and heat capacity at temperatures up to

1500oC, were found to have a great deal of influence on the shape of simulated time-

temperature curves. It should be mentioned, however, that these apparent thermal

properties, to some degree, implicitly account for the physical phenomena other than

heat transfer, such as mass transfer and phase change. This happens because the

temperature rise in LSF walls exposed to fire is affected by processes not described by

heat transfer, such as ablation of plasterboard, migration of moisture vapours, and

penetration of cool ambient air or hot furnace gases into the cavity. Figures 3.31 to 3.34

show the idealized thermal properties of glass fibre and rock fibre.

Measure specific heat and mass loss of rock and glass fibre insulations

Use specific heat and mass loss from DSC test results and thermal conductivity from

previous research work to perform finite element modelling using SAFIR. For

gypsum plasterboard, use idealised thermal properties that was obtained earlier using

the procedure described in Chart 3.1

Compare small scale test experimental results

from Kolarkar (2010) with numerical results

(Specimens 6 to 11)

Compare load bearing wall experimental

results from Kolarkar (2010) with numerical

results (Tests 4 to 7)

Idealised thermal properties of rock and glass fibre insulations

Revise thermal properties

Re-run SAFIR

Perform DSC tests of rock and glass fibre insulations



Figure 3.27: Specific Heat of Plasterboard Reported by Various Researchers

Figure 3.28: Proposed Specific Heat of Plasterboard

0

5000

10000

15000

20000

0 200 400 600 800 1000 1200

Spe

cifi

c H

eat

(J

/kg.

oC

)

Temperature (oC)

Manzello (2006) Sultan (1996) Thomas (2010)

-5000

0

5000

10000

15000

20000

0 200 400 600 800 1000 1200

Spe

cifi

c H

eat

(J

/kg.

oC

)

Temperature (oC)

SAFIR Idealised Experimental



Table 3.3: Proposed Specific Heat of Plasterboard

Temperature (oC) Specific Heat (J/kg.

oC)

0 950

20 950

100 950

140 17500

156 12500

170 13000

200 950

660 950

670 3000

680 950

1200 950

4000 950

950pC CTC oo 200

950pC CTC oo 10020

4042575.413 TCp CTC oo 140100

612505.312 TCp CTC oo 156140

6.6928471.35 TCp CTC oo 170156

8128367.401 TCp CTC oo 200170 (3.4)

950pC CTC oo 660200

134350205 TCp CTC oo 670660

140350205 TCp CTC oo 680670

950pC CTC oo 4000680

950pC For other temperatures



Figure 3.29: Proposed Thermal Conductivity of Plasterboard

Table 3.4: Proposed Thermal Conductivity of Plasterboard

Temperature (oC) Thermal Conductivity (W/m.

oC)

0 0.25

140 0.25

150 0.13

300 0.13

800 0.18

1200 0.30

1201 0.80

4000 10.0

25.0k CTC oo 1400

Tk 012.093.1 CTC oo 150140

13.0k CTC oo 300150 (3.5)

Tk 0001.01.0 CTC oo 800300

0.00

0.20

0.40

0.60

0.80

1.00

0 500 1000 1500

The

rmal

Co

nd

uct

ivit

y (W

/m.o

C)

Temperature (oC)

SAFIR (2004) Rahmanian (2009)

Thomas (2002) Idealised Thermal Conductivity



06.00003.0 Tk CTC oo 1200800

1633.30033.0 Tk CTC oo 40001201

Figure 3.30: Proposed Relative Density of Plasterboard

Table 3.5: Proposed Relative Density of Plasterboard

Temperature (oC) Density (kg/m

3) Relative Density (%)

0 729 100

120 729 100

170 656 90

1500 656 90

0.1RD CTC oo 1200

TRD 002.024.1 CTC oo 170120 (3.6)

9.0RD CTC oo 1000170

0

10

20

30

40

50

60

70

80

90

100

110

0 100 200 300 400 500

Re

lati

ve D

en

sity

(%

)

Temperature (oC)

Plasterboard 1 Plasterboard 2 Plasterboard 3

Plasterboard 4 Proposed Density



Figure 3.31: Specific Heat of Rock Fibre Insulation

Figure 3.32: Thermal Conductivity of Rock Fibre Insulation

0

1000

2000

3000

4000

5000

50 100 150 200 250

Spe

cifi

c H

eat

(J/

kg.o

C)

Temperature (oC)

Rock Fibre 1 Rock Fibre 2 Apparent Specific Heat

0.00

0.50

1.00

1.50

2.00

2.50

3.00

3.50

4.00

0 200 400 600 800 1000 1200

The

rmal

Co

nd

uct

ivit

y (

W/m

.oC

)

Temperature (oC)

Takeda (2001) Thomas (2002)

Alfawakhiri (2001) Idealised Thermal Conductivity



Table 3.6: Proposed Thermal Conductivity of Rock Fibre Insulation


oC)

0 0.25

550 0.30

1200 2.00

Tk 00009.025.0 CTC oo 5500

Tk 0026.01385.1 CTC oo 1200550 (3.7)

Figure 3.33: Specific Heat of Glass Fibre Insulation

0

1000

2000

3000

4000

5000

50 100 150 200 250 300 350 400 450 500

Spe

cifi

c H

eat

(J/

kg.o

C)

Temperature (oC)

Glass Fibre 1 Glass Fibre 2 Apparent Specific Heat



Figure 3.34: Thermal Conductivity of Glass Fibre Insulation

Table 3.7: Proposed Thermal Conductivity of Glass Fibre Insulation


oC)

0 0.5

600 0.6

700 2.0

800 10000

Tk 0002.05.0 CTC oo 6000

Tk 014.08.7 CTC oo 700600 (3.8)

6998498.99 Tk CTC oo 800700

0.00

0.50

1.00

1.50

2.00

2.50

3.00

3.50

4.00

0 200 400 600 800 1000 1200

The

rmal

Co

nd

uct

ivit

y (

W/m

.oC

)

Temperature (oC)

Takeda (2001) Alfawakhiri (2001) Idealised Thermal Conductivity



3.6. Summary

This chapter has presented the details of an experimental study to determine the thermal

properties of the materials used in LSF wall panels. Properties of interest include

specific heat, density and thermal conductivity as a function of temperature. Eight DSC

tests were conducted to measure specific heat and mass loss of the materials. They

included four tests for BORAL Firestop gypsum plasterboard, two tests for BORAL

glass fibre insulation and two tests for Tombo Brand M.G. Mighty Rock Fibre.

The values produced from the DSC tests were then converted to specific heat by using

the calculation method provided in the SETARAM TGA DSC manual. Thermal

property values from the DSC tests and other previous works were calibrated to

accurately predict the time-temperature profiles of plasterboards and their assemblies

including LSF walls using finite element models in comparison with Kolarkar’s (2010)

experimental results.

The idealised thermal property data set provides valuable information that can be used

to model the thermal behaviour of LSF wall panels under fire conditions.

Finite Element Analyses of Small Scale Plasterboard Panels


Chapter 4


4.1. General

In order to understand the thermal performance of LSF wall panels made of gypsum

board, glass fibre and rock fibre insulations, eleven small scale tests were conducted in

the Fire Research Laboratory of Queensland University of Technology. Thermal

performance of single, double and triple layers of plasterboards was studied. Different

types of insulations were also used to help improve their fire performance. Composite

panels were also developed with a layer of insulation between two sheets of

plasterboard.

Recently many numerical heat transfer models have been developed (Franssen et al.,

2005, Sultan et al., 1996). There are also many general finite element packages that can

be used for thermal analyses. The computational model employed in this study to

predict the thermal behaviour of the tested LSF wall panels assemblies was SAFIR2007.

This chapter presents the capabilities and limitations of SAFIR2007 and also the details

of the development of finite element models to simulate the behaviour of tested

plasterboard panels using SAFIR and GID pre and post processors. It also presents the

details of finite element modelling of small scale tests performed on individual and

multiple layers of plasterboard. It also examines the thermal performance of composite

panels developed from different insulating materials of varying densities and thickness.

The thermal properties of materials used in finite element modelling were obtained

based on calibrating the numerical results against the results from these 11 small scale

tests (Chapter 3).

4.2. SAFIR

SAFIR is a computer software developed at the University of Liege for the simulation

of the behaviour of building structures subjected to fire. The fire is introduced as a data



(in terms of a curve giving either the evolution of the gas temperature in the fire

compartment or the evolution of the net flux on the surface of the structure) and the

software calculates the evolution of the temperature in the structural elements which can

be discretized in 2D or 3D. It is viewed today as the second generation of structural fire

codes developed in Liege, the first generation being a computer program called

Computer Engineering of the Fire Design of Composite and Steel Structures

(CEFICOSS).

In SAFIR 2007, heat transfer by convection and radiation are modelled at the boundary

conditions. Internal cavities are permitted in two-dimensional analysis, with radiation

and convection modelled along void boundaries. SAFIR 98 only allowed heat transfer

to be modelled through a completely closed void, totally surrounded by boundary

elements. However, further modifications to the program by Franssen et al. (2007) led

to the ability to model heat transfer across an open void defined by an axis of symmetry

at the cavity opening. This feature is now incorporated in SAFIR 2007 (Franssen et al.,

2007).

Below is a summary describing the functionalities of the software SAFIR and what it

can do:

-. The temperature distribution is transient; it varies as a function of time.

-. The basic equation for conduction in the elements is the Fourier equation.

-. 2D or 3D calculations can be performed.

-. The finite elements have a temperature distribution that varies linearly along the

borders of the elements.

-. The elements are triangular or quadrilateral (not necessarily regular) for the 2D

calculations. The elements are prismatic (not necessarily regular) with 6 or 8

nodes for the 3D calculations.

-. The geometry of the elements does not change during the calculation (no

prediction of spalling).



-. The mechanical behaviour of the structure does not influence the temperature

distribution.

-. One single material is present in each element. Different materials can be

present in different elements.

-. There is no contact resistance between adjacent elements.

-. Some predefined thermal material models are embedded in the code, namely

concrete, steel, wood, gypsum and aluminium materials. Among these thermal

properties, the thermal conductivity, the specific heat, and mass are temperature

dependent. The relative emissivity and the coefficient of conductivity are

constants.

-. The energy required to evaporate eventual liquid water is taken into account.

The energy required to heat the liquid water or the vapour is neglected.

-. Boundary conditions are either adiabatic (axes of symmetry), or a prescribed

temperature-time curve in the ambiance plus heat flux calculated from linear

convection and radiation, or an imposed heat flux.

-. Some prescribed temperature – time curves are embedded in the code such as

ISO834 and the ASTM E119 curves.

-. Radiation in the internal cavities can be taken into account in 2D calculation

only. The air is supposed to be transparent. Convection in the cavity is taken into

account in an approximate manner.

4.3. Limitations of SAFIR

Although SAFIR is a very powerful finite element program, the program deficiencies

and limitations exist in its ability to model complex wall assemblies. These limitations

are discussed in this section.



4.3.1. Moisture Movement

Heat transfer within gypsum is highly dependent on the moisture content of the

material. The user has the ability to account for moisture content within the material by

modifying the respective specific heat curve in the model. However, modelling moisture

movement across the cavity is a more complex problem, which is not considered in

SAFIR. This phenomenon is generally neglected due to its complexity, and because it

only influences the heat transfer across the cavity at temperatures below 120oC.

4.3.2. Ablation

Ablation is the process by which consecutive thin layers of gypsum are shed from the

lining. This has the effect of reducing the cross-sectional thickness of the gypsum

lining, thus increasing the heat flux across the lining. SAFIR does not allow the user to

remove elements from the section to simulate ablation, and therefore, it must be taken

into account when defining the thermal properties of the lining.

4.3.3. Shrinkage

Shrinkage and cracking of the lining are typically taken into account by increasing the

thermal conductivity of the lining once dehydration has occurred. However, another

phenomenon occurs within the assembly due to moisture movement within materials.

This creates a void between the lining and the stud, altering the form of heat transfer

into the stud. SAFIR does not currently allow the user to modify the dimension of

elements during calculation. A possible way of accounting for this effect is to pre-define

a gap between the stud and lining, assigning it with the properties of initial material, and

then alter the properties to that of air once shrinkage is expected to have occurred.

Further investigation of this phenomenon and its influence on modelling results will be

left for recommended study.



4.4. GID Pre and Post Processor

GID is a universal pre and post processor for numerical simulations in science and

engineering. It has been designed to cover all the common needs in the numerical

simulations field from pre to post processing such as geometrical modelling, effective

definition of analysis data, mesh generation, transfer data to analysis software and

visualisation of results. In this study GID software was used to create the input file for

finite element modelling as well as analysing the model output result. Below are the

steps that were used to construct the model and how to use the output results.

4.4.1. SAFIR Problem Types

Within GID, all SAFIR input parameters, standard materials and sections are made

available through the selection of a SAFIR ‘problem type’ (see Figure 4.1). Four

individual problem types may be chosen corresponding to two-dimensional thermal

analysis, two-dimensional structural analysis, three-dimensional thermal analysis and

three-dimensional structural analysis. In this study ‘Safir_Thermal_2d’ problem type

was used.

Figure 4.1: SAFIR Problem Types

4.4.2. Model Geometry

In terms of SAFIR, the GID y-coordinate equals the SAFIR y-coordinate (the first

global coordinate), the GID x-coordinate is the second global coordinate, which is

denoted as z-coordinate in SAFIR. The geometrical model may be input into GID



manually or using Computer Aided Drawing (CAD) software via direct import of DXF

drawing file. Figure 4.2 shows two GID geometries that were used in this study.

(a) Small Scale Test Specimen 9 Geometry

(b) Load Bearing Wall Typical Geometry

Figure 4.2: Typical GID Geometry

4.4.3. Materials

All materials embedded in SAFIR may be applied to NURBS surface (magenta colour

line in Figure 4.2) within GID. Properties of user defined materials may also be input

and applied to NURBS surfaces in a similar fashion. Figure 4.3 shows the GID interface

with the material conditions while Figure 4.4 shows the small scale Test Specimen 6

with materials.

(a) Material Embedded in SAFIR (b) User Defined Material

Figure 4.3: GID Interface for Material Condition



Figure 4.4: Small Scale Test Specimen 6 with Material

The emissivity of the exposed gypsum board should be dependent on the state of the

thermal degradation of its surface (Clancy 1999). In SAFIR a relative emissivity

coefficient is used to represent the surface emissivity of the board at all temperatures. A

similar approach was adopted for the coefficient of convection for both the cold and hot

surfaces. The thermal properties that were used in this study are:

-. Convection Coefficient Hot Surface = 25 for all materials.

-. Convection Coefficient Cold Surface = 10 for all materials.

-. Relative Emissivity = 0.9 for plasterboard and insulation and 0.6 for steel.

-. Thermal Conductivity (W/m.K) – Refer to Chapter 3.

-. Specific Heat (J/kg.K) – Refer to Chapter 3.

-. Specific Mass (kg/m3) – Refer to Chapter 3.

4.4.4. Boundary Conditions

SAFIR provided some predefined temperature curves such as FISO, F20, F1000, F0,

etc. These entire predefined temperature curves can be applied directly to a point or a

line in the model geometry. User defined temperature can also be applied in a similar

manner. FISO was used in the line where the model was exposed to standard fire curve

produced by the furnace while F20 was used in the ambient side (see Figure 4.5).

Figure 4.5: Specimen 3 Boundary Conditions



F20 = Temperature at 20oC

FISO = Standard Time-Temperature curve according to AS 1530.4



4.4.5. MESHING

GID can create either triangular or quadrilateral meshes for the 2D calculations. For 3D

calculations prismatic with 6 or 8 nodes is normally used. GID displays a dialog box

where element size can be entered which is used in the case of non-structured mesh (see

Figure 4.6). It will display the number of nodes and elements it created (see Figure 4.7).

Figure 4.8 shows some typical small scale and load bearing wall model with mesh

generated.

Figure 4.6: Mesh Generation Dialog Box

Figure 4.7: Summary of Mesh Generated



(a) Specimen 8 Mesh

(b) Load Bearing Wall with Glass Fibre Cavity Insulation

Figure 4.8: Generated Finite Element Mesh

4.4.6. General Data

GID displays a dialog box where general data for SAFIR calculation can be entered (see

Figure 4.9). All variables have the same name as in the SAFIR reference manual and

have predefined values.

Figure 4.9: SAFIR Problem Data



where:

TETA : Parameter for the time integration

TINITIAL : Temperature of the structure at time t=0

Global Centre (Y0, Z0) : First and second global coordinate of the centre of the

cross section for the structural calculation

Centre of Torsion (Yc, Zc) : Coordinate of the centre of torsion

NVOID : The number of voids in the cross section

NFRONTIERVOID : Maximum number of surfaces enclosing the internal

void

TIMESTEP : Time step in seconds

UPTIME : End time in seconds

TIMEPRINT : Time step for printing results

4.4.7. Post Processing

GID can be used as a post-processor to graphically plot the results contained in the

SAFIR analysis output file. In the post-processing mode GID is capable of displaying

thermal contours, plotting the temperature history of identified node/element and for a

structural analysis displaying resulting load vectors and structural actions. Figure 4.10

shows the GID with active post-processing interface and temperature contours.

Figure 4.10: GID Post-Process Interface with Temperature Contours Active



4.5. Model Configuration

Eleven fire tests were conducted on small scale specimens of sizes 1350 mm x 1080

mm. Figure 4.11 shows the test set-up of small scale tests. The specimen was exposed

to heat from one side only and the time-temperature profiles at various locations across

the thickness of the test specimens were measured by using metal sheathed

thermocouples to help assess their fire performance. Table 4.1 shows the positions of

thermocouples as indicated by the coloured dots (Kolarkar, 2010). The temperature rise

of these thermocouples served as the input to the computer controlling the furnace heat

according to the cellulosic fire curve (Standard time-temperature curve) given in AS

1530.4 (SA, 2005), which is similar to ISO 834-1 (1999) and ASTM E119 (1995). Tests

were stopped once the plasterboard paper on the ambient side of the specimen started to

burn.

(a) Large Furnace (b) Single Burner for Small Scale Test

Figure 4.11: Test Set-up of Gypsum Plasterboard (Kolarkar, 2010)



Table 4.1: Details of Plasterboard Test Specimens

No.

Configuration

Specimen Description

1

Pb = 13 mm

2

Pb = 16 mm

3

Pb1 = 13 mm (Fire Side)

Pb2 = 16 mm (Ambient Side)

4



5


Pb2 = 16 mm (Central)


6,7,8,9


Insulation = Glass Fibre of varying thickness,

density and type


10,11


Insulation = Rock Fibre of varying thickness,

density and type


4.6. Small Scale Test Specimen 1

Test Specimen 1 was made of a single layer of 13 mm thick BORAL Firestop gypsum

plasterboard with 729 kg/m3 in density. Thermocouples were located on the specimen as

shown in Table 4.1. One side of Test Specimen 1 was subjected to the standard time-

temperature heating regime in the furnace (see Figure 4.12(a)). By the end of 3 minutes

smoke was seen to start coming from the edges of the specimen. This was on account of

the burning of the plasterboard paper on the exposed side. The smoke subsided after the

paper was completely burnt. By the end of 6 minutes steam was seen to come out from

the specimen and condense on the top front face of the furnace adapter. By the end of 12



to 13 minutes the steam subsided and the specimen was soon seen to burn steadily

without letting out smoke or steam. By the end of 18 minutes, the ambient side paper of

the plasterboard started to discolour. The specimen was also seen to bow laterally

outward (see Figure 4.12(b)). By the end of 33 minutes the outside paper had started to

burn and the test was stopped. This was considered to be the failure point. At the end of

the test the ambient side temperature reached 266oC.

(a) Specimen 1 at the start of the test (b) Specimen 1 Thermal Bowing

Figure 4.12: Small Scale Test Specimen 1

Figure 4.13 shows the time-temperature profiles across the plasterboard thickness for

Test Specimen 1 and compares them with the results from finite element modelling.

Figure 4.14 shows the temperature distributions in the cross-section of plasterboard.

Figure 4.13: Time - Temperature Profiles of Test Specimen 1

(13 mm Plasterboard) from Experiment and FEA

0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 5 10 15 20 25 30

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - 7 mm From Fire Side

Experiment - Ambient Side SAFIR - Fire Side

SAFIR - 7 mm From Fire Side SAFIR - Ambient Side



(a) 1 Minute

(b) 15 Minutes

(c) 30 Minutes

Figure 4.14: Specimen 1 Temperature Distributions from FEA

The validation of the developed finite element model was achieved by comparing the

time-temperature profiles of the plasterboard from FEA with test results. The predicted

fire side temperatures are in excellent agreement with test results. At all temperature

measuring locations, the correlation between numerical and test results is quite good but

is not exact. However, considering software limitations, the agreement is reasonable.

Figure 4.13 shows that the model developed to predict the time-temperature profiles

give good accuracy. Table 4.2 results confirm this with an overall mean of 0.986 and an

overall coefficient of variation of 0.152.



Table 4.2: Comparison of Experimental and Finite Element Analysis Results for

Test Specimen 1

Time Fire Side 7mm from Fire Side Ambient

Exp FEA FEA/EXP Exp FEA FEA/EXP Exp FEA FEA/EXP

0 95 20 0.21 30 20 0.67 22 20 0.91

1 143 136 0.95 60 38 0.63 24 21 0.89

2 317 186 0.59 85 55 0.65 33 37 1.10

3 418 346 0.83 95 61 0.64 67 45 0.68

4 471 434 0.92 100 83 0.83 79 54 0.68

5 509 485 0.95 110 104 0.95 80 66 0.82

6 551 520 0.94 115 120 1.04 89 72 0.81

7 579 549 0.95 125 136 1.09 90 75 0.84

8 603 579 0.96 135 166 1.23 91 78 0.86

9 622 604 0.97 150 191 1.27 92 86 0.93

10 642 627 0.98 160 219 1.37 101 91 0.90

11 658 645 0.98 195 239 1.23 104 95 0.91

12 672 661 0.98 215 257 1.20 107 98 0.92

13 688 675 0.98 225 278 1.24 109 99 0.91

14 702 687 0.98 240 296 1.23 111 104 0.94

15 714 699 0.98 270 320 1.19 130 108 0.83

16 725 710 0.98 285 348 1.22 140 117 0.84

17 734 722 0.98 300 377 1.26 149 124 0.83

18 742 733 0.99 320 401 1.25 166 139 0.84

19 750 742 0.99 350 421 1.20 191 171 0.90

20 757 752 0.99 380 453 1.19 216 216 1.00

21 764 761 1.00 405 476 1.18 237 232 0.98

22 771 769 1.00 420 491 1.17 240 242 1.01

23 778 777 1.00 425 502 1.18 250 248 0.99

24 784 784 1.00 430 509 1.18 253 252 1.00

25 790 790 1.00 435 516 1.19 257 255 0.99

26 796 796 1.00 440 521 1.18 260 258 0.99

27 801 802 1.00 445 526 1.18 263 260 0.99

28 807 808 1.00 445 530 1.19 266 262 0.98

29 811 813 1.00 450 533 1.18 266 263 0.99

30 817 818 1.00 450 537 1.19 266 265 1.00

Mean 0.938 1.110 0.911

StDev 0.156 0.205 0.095

CoV 0.166 0.184 0.105

Overall Mean 0.986

Overall CoV 0.152




Test Specimen 2 was made of a single layer of 16 mm thick BORAL Firestop gypsum

plasterboard with 729 kg/m3 density. Thermocouples were located on the specimen as

shown in Table 4.1. The specimen was fire tested for about 78 minutes. One side of Test

Specimen 1 was subjected to the standard time-temperature heating regime in the

furnace (see Figure 4.15 (a)). The observations pertaining to the evolution of smoke and

steam were similar to that of test Specimen 1. By the end of 29 minutes, the paper on

the ambient surface started to discolour uniformly. By 40 minutes, the ambient surface

had become quite dark. Towards the end of the test, the paper was partially burnt and

the specimen had begun to bow laterally in the outward direction (see Figure 4.15(b)).

By 78 minutes, the ambient surface temperature reached 271oC.

(a) Specimen 2 at the Start of the Test (b) Specimen 2 at the End of the Test








(16 mm Plasterboard) from Experiment and FEA

(a) 1 Minute

(b) 39 Minutes


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 10 20 30 40 50 60 70 80

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - 12 mm Experiment - 8 mm

Experiment - 4 mm Experiment - Ambient SAFIR - Fire Side

SAFIR - 12 mm SAFIR - 8 mm SAFIR - 4 mm

SAFIR - Ambient



(c) 78 Minutes


Specimen 1 failed at 30 minutes with the ambient side temperature reaching 266oC

while Specimen 2 failed at 78 minutes with the ambient side temperature reaching

271oC. The failure was considered to be the point at which the ambient side paper

started burning. The observed temperatures were found to be much higher than those

given in AS 1530.4 for insulation criterion (140oC average and 180

oC maximum). It can

be observed that 3mm (25%) increase in thickness from Specimen 1 to Specimen 2

doubles the fire rating of the gypsum plasterboard.


Test Specimen 2

FS 4mm FS 8mm FS 12mm FS Amb

Mean 0.983 1.208 1.160 1.146 0.968

StDev 0.063 0.159 0.148 0.177 0.118

CoV 0.064 0.131 0.128 0.155 0.121

Overall Mean 1.093

Overall CoV 0.120





is not exact. Figure 4.16 shows that the model developed to predict the time-temperature



profiles give good accuracy. Table 4.3 results confirm this with an overall mean of

1.093 and an overall coefficient of variation of 0.120.


Test Specimen 3 was made of one layer each of 13 mm and 16 mm thick BORAL

Firestop gypsum plasterboard with 729 kg/m3 in density. The 13 mm thick plasterboard

was labelled as Pb1 and formed the exposed side of the specimen, whereas the 16 mm

thick plasterboard was labelled as Pb2 and formed the ambient side of the specimen.

Thermocouples were located on the specimen as shown in Table 4.1. The specimen was

fire tested for about 171 minutes (see Figure 4.18(a)). The fire side paper of the exposed

plasterboard caught fire by the end of 3 minutes when the temperature of the exposed

surface was about 400oC. The smoke was soon followed by a period of steady burning

during which time there was hardly any emission of smoke or steam. By the end of 20

minutes, smoke reappeared. This was probably due to the plasterboard paper on the

ambient side of Pb2 (Plasterboard 2) burning. By the end of 62 minutes the plasterboard

on the unexposed surface of the specimen started to discolour, when its temperature was

about 200oC. Towards the end of the test, the paper on the ambient surface of the

specimen was noticed to have blackened uniformly (see Figure 4.18(b))









(13 & 16 mm Plasterboards) from Experiment and FEA


Test Specimen 3


Mean 0.967 1.055 1.014 1.203 0.979

StDev 0.080 0.055 0.152 0.168 0.116

CoV 0.082 0.052 0.150 0.140 0.118

Overall Mean 1.044

Overall CoV 0.109

0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 20 40 60 80 100 120 140 160 180

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - 7 mm From FS

Experiment - 13 mm From FS Experiment - 21 mm From FS

Experiment - Ambient Side SAFIR - Fire Side

SAFIR - 7 mm From FS SAFIR - 13 mm From FS

SAFIR - 21 mm From FS SAFIR - Ambient Side



(a) 1 Minute

(b) 86 Minutes

(c) 171 Minutes


Specimen 3 failed at 171 minutes with the ambient side temperature reaching 249oC. In

terms of thermal performance, Specimen 3 is almost six times better than Specimen 1

and two times better than Specimen 2. It is to be noted that Specimen 3 is 2.25 times

thicker than Specimen 1 and 1.8 thicker than Specimen 2.











Test Specimen 4 was made of two layers of 16 mm thick BORAL Firestop gypsum

plasterboard with 729 kg/m3 in density. The first 16 mm thick plasterboard was labelled

as Pb1 and formed the exposed side of the specimen, whereas the second 16 mm thick

plasterboard was labelled as Pb2 and formed the ambient side of the specimen.

Thermocouples were located on the specimen as shown in Table 4.1. The specimen was

fire tested for about 222 minutes (see Figure 4.21(a)). The behaviour of Specimen 4 was

very much similar to that of Specimen 3 test. After intermittent evolution of smoke and

steam, the ambient side of the specimen started to discolour at the end of 78 minutes.

The test was continued for some time even after burning of the ambient side paper. The

specimen displayed a small amount of lateral deflection in the outward direction. The

test was finally stopped when most of the ambient side paper started to peel and burn

(see Figure 4.21(b)).

Specimen 4 failed at 222 minutes with the ambient side temperature reaching 249oC. Its

failure time is 7 times longer compared to Specimen 1 although it is only 2.5 times

thicker. It means an increase of 19 mm thick gypsum plasterboard gives an additional

192 minutes increase in failure time. It is on average of 10 minutes per 1 mm thick

gypsum plasterboard. This prediction seems to match well with the failure times of

Specimens 2 and 3. Specimen 2 has an additional 3 mm thickness which gives about 30

minutes more in failure time. Specimen 3 has an additional 16 mm thickness which

gives about 160 minutes more in failure time.






Test Specimen 4


Mean 0.973 1.068 1.008 1.137 1.006

StDev 0.071 0.056 0.099 0.129 0.132

CoV 0.073 0.053 0.098 0.113 0.131

Overall Mean 1.038

Overall CoV 0.094





time-temperature profiles of the plasterboard from FEA with test results. At all

temperature measuring locations, the correlation between numerical and test results is

quite good but is not exact. Figure 4.22 shows that the model developed to predict the

time-temperature profiles give good accuracy. Table 4.5 results confirm this with an

overall mean of 1.038 and an overall coefficient of variation of 0.094.




(Two 16 mm Plasterboards) from Experiment and FEA

(a) 1 Minute

(b) 111 Minutes


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 50 100 150 200

Tem

pe

ratu

re (

oC

)

Time (min)


Experiment - 16 mm From FS Experiment - 24 mm From FS

Experiment - Ambient SAFIR - Fire Side

SAFIR - 8 mm From FS SAFIR - 16 mm From FS

SAFIR - 24 mm From FS SAFIR - Ambient



(c) 222 Minutes



Test Specimen 5 was made of three layers of 16 mm thick BORAL Firestop gypsum

plasterboard with 729 kg/m3 in density. The first 16 mm thick plasterboard was labelled

as Pb1 and formed the exposed side of the specimen, the second plasterboard was

sandwiched between plasterboard 1 and plasterboard 3 and was labelled as Pb2,

whereas the third 16 mm thick plasterboard was labelled as Pb3 and formed the ambient

side of the specimen. Thermocouples were located on the specimen as shown in Table

4.1. The specimen was fire tested for about 186 minutes. The behaviour of Specimen 5

was very much similar to that of Specimen 4 test. Plasterboard 1 was seen to heat up

quite rapidly with its temperature reaching 900oC by the end of 155 minutes. At about

165 minutes Plasterboard 1 must have partially or fully collapsed as the curve is seen to

rise rapidly and merge with the fire side (FS) curve. At the end of the test, the

temperature across the Pb2-Pb3 interface had reached 750oC and the unexposed surface

had crossed 200oC.

Based on earlier assumption, Specimen 5 should give an additional 350 minutes more

than Specimen 1 failure time and it should fail at 380 minutes. But in this test, it failed

at 186 minutes. It even performed worse than Specimen 4 even though it had an

additional 16 mm thick gypsum plasterboard. This may be because of the fluctuation in

the furnace at around 140 minutes (see Figure 4.24). Because of this sudden fluctuation

in the furnace temperature the time-temperature profile at the Fire Side (FS), 16mm

from FS, 32mm from FS and Ambient Side suddenly increased.




(Three 16 mm Plasterboards) from Experiment and FEA

(a) 1 Minute

(b) 93 Minutes


0

200

400

600

800

1000

1200

1400

-10 10 30 50 70 90 110 130 150 170 190

Tem

pe

ratu

re (

oC

)

Time (min)


Experiment - 32 mm From FS Experiment - Ambient

SAFIR - Fire Side SAFIR - 16 mm From FS

SAFIR - 32 mm From FS SAFIR - Ambient



(c) 186 Minutes






Test Specimen 5

FS 16mm FS 32mm FS Amb

Mean 0.964 0.965 0.891 0.831

StDev 0.087 0.113 0.124 0.123

CoV 0.090 0.117 0.139 0.148

Overall Mean 0.913

Overall CoV 0.124











Test Specimen 6 consisted of a composite panel formed by sandwiching a layer of glass

fibre insulation between the plasterboards. This specimen has a cavity of 32 mm that

was filled with glass fibre. Boral Firestop plasterboard with 729 kg/m3 in density was

used while the density of the glass fibre is 21.68 kg/m3. Thermocouples were located on

the specimen as shown in Table 4.1. The specimen was fire tested for about 181

minutes. The initial behaviour of this specimen was similar to the previously tested

specimens. The ambient surface of this specimen showed uniform discolouration after

about 110 minutes of testing (see Figure 4.26(a)). The test was stopped when the paper

on the ambient surface started to burn. When the Specimen 6 was inspected after the

test, it was noted that the glass fibre insulation has been almost completely consumed by

the heat with only small amounts still visible along the edges of the specimen (see

Figure 4.26(b)). Figure 4.27 shows the time-temperature profiles across the plasterboard

thickness for Test Specimen 6 and compares them with the results from finite element

modelling.

(a) Specimen 6 near the End of the Test (b) Glass Fibre Consumed by Heat



Test Specimen 6

FS Pb1 - Ins Ins - Pb2 Amb

Mean 0.968 0.949 0.947 0.971

StDev 0.084 0.156 0.191 0.162

CoV 0.086 0.164 0.202 0.167

Overall Mean 0.959

Overall CoV 0.155



Figure 4.27: Time - Temperature Profiles of Test Specimen 6 from Experiment

and FEA

Figure 4.28 shows the temperature distributions in the cross-section of Specimen 6.

(a) 1 Minute


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 20 40 60 80 100 120 140 160 180

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Insulation

Experiment - Insulation - Pb2 Experiment - Ambient

SAFIR - Fire Side SAFIR - Pb1 - Insulation

SAFIR - Insulation - Pb2 SAFIR - Ambient



(b) 91 Minutes

(c) 181 Minutes



temperature profiles of the plasterboard from FEA with test results. The predicted fire

side temperatures are in excellent agreement with test results. At all temperature


is not exact. Figure 4.27 shows that the model developed to predict the time temperature







used while the density of the glass fibre is 43.4 kg/m3. Thermocouples were located on


minutes. The behaviour of this specimen was similar to that of Specimen 6. The test was

stopped when the paper on the ambient surface started to burn. When Specimen 7 was



inspected after the test, it was noted that the glass fibre insulation has been almost

completely consumed by the heat with only small amounts still visible along the edges

of the specimen. Figure 4.29 shows the time-temperature profiles across the

plasterboard thickness for Test Specimen 7 and compares them with the results from

finite element modelling. Figure 4.30 shows the temperature distributions in the cross-

section of Specimen 7.


and FEA

(a) 1 Minute


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 20 40 60 80 100 120 140 160 180

Tem

pe

ratu

re (

oC

)

Time (min)







(b) 89 Minutes

(c) 179 Minutes



temperature profiles of the plasterboard from FEA with test results. At all temperature






Test Specimen 7


Mean 0.977 0.963 1.077 1.023

StDev 0.062 0.149 0.178 0.157

CoV 0.063 0.155 0.166 0.154

Overall Mean 1.010

Overall CoV 0.134







used while the density of the glass fibre is 37 kg/m3. Thermocouples were located on


minutes. The behaviour of this specimen was similar to that of Specimens 6 and 7. The

test was stopped when the paper on the ambient surface started to burn (see Figure

4.31(a)). When Specimen 8 was inspected after the test, it was noted that the glass fibre

insulation has been almost completely consumed by the heat with only small amounts

still visible along the edges of the specimen (see Figure 4.31(b)).

(a) Specimen 8 near the End of the Test (b) Glass Fibre Consumed by Heat




Figure 4.33 shows the temperature distributions in the cross-section of Specimen 8.

(a) 1 Minute




(b) 100 Minutes

(c) 200 Minutes



and FEA

0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)














Test Specimen 8


Mean 0.982 1.041 1.167 1.149

StDev 0.054 0.173 0.205 0.212

CoV 0.055 0.166 0.176 0.185

Overall Mean 1.085

Overall CoV 0.145



fibre insulation between the plasterboard. This specimen has a cavity of 13 mm that was

filled with glass fibre. Boral Firestop plasterboard with 729 kg/m3 in density was used

while the density of the glass fibre is 168 kg/m3. Thermocouples were located on the

specimen as shown in Table 4.1. The specimen was fire tested for about 185 minutes.

The behaviour of this specimen was similar to that of Specimens 6, 7, and 8. The test

was stopped when the paper on the ambient surface started to burn. When Specimen 9

was inspected after the test, it was noted that the glass fibre insulation has been almost

completely consumed by the heat with only small amounts still visible along the edges

of the specimen. Figure 4.34 shows the time-temperature profiles across the

plasterboard thickness for Test Specimen 9 and compares them with the results from

finite element modelling. Figure 4.35 shows the temperature distributions in the cross-

section of Specimen 9.




and FEA

(a) 1 Minute

(b) 93 Minutes


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 20 40 60 80 100 120 140 160 180

Tem

pe

ratu

re (

oC

)

Time (min)







(c) 185 Minutes



temperature profiles of the plasterboard from FEA with test results. There was a drop in

the furnace temperature around 110 minutes, which caused some problems in relation to

accurately predicting the time-temperature profiles. Overall the correlation between

numerical and test results is quite good but is not exact. Figure 4.34 shows that the

model developed to predict the time temperature profiles give good accuracy. Table

4.10 results confirm this with an overall mean of 0.977 and an overall coefficient of

variation of 0.174.


Test Specimen 9


Mean 0.978 0.838 1.131 0.963

StDev 0.057 0.093 0.356 0.203

CoV 0.058 0.111 0.315 0.210

Overall Mean 0.977

Overall CoV 0.174


Test Specimen 10 consisted of a composite panel formed by sandwiching a layer of

rock fibre insulation between the plasterboards. This specimen has a cavity of 25 mm



that was filled with rock fibre. Boral Firestop plasterboard with 729 kg/m3 in density

was used while the density of the rock fibre was 100 kg/m3. Thermocouples were

located on the specimen as shown in Table 4.1. The specimen was fire tested for about

174 minutes (see Figure 4.36(a)).

(a) Specimen 10 at the Start of the Test (b) Rock Fibre after Test


Slight thermal bowing in the outward direction was noted at the end of the test. The

Plasterboard 1 – Insulation temperature profile was seen to rise rapidly to about 600oC

by the end of 32 minutes, beyond which it flattened out, with the temperature gradually

increasing to 900oC by the end of 147 minutes. Around this time, plasterboard 1 must

have collapsed as the curve jumped rapidly to merge with the Fire Side curve. The

failure of Plasterboard 1 affecting Insulation – Plasterboard 2 and Ambient Side can be

seen by the sudden increase in both temperature profiles (see Figure 4.37). Contrary to

glass fibre insulation, the rock wool insulation showed greater resistance to

disintegration. The physical presence of the insulation was blocking and redirecting the

heat flow back to Plasterboard 1. This resulted in the rising of temperature of

Plasterboard 1 – Insulation to values beyond 700oC and steadily kept rising up to 900

oC

when Plasterboard 1 started to breach. Even after getting directly exposed to fire after

the collapse of Plasterboard 1, the insulation remained intact and continued to offer

protection to Plasterboard 2 (see Figure 4.36(b)). Figure 4.37 shows the time-

temperature profiles across the plasterboard thickness for Test Specimen 10 and



compares them with the results from finite element modelling. Figure 4.38 shows the

temperature distributions in the cross-section of Specimen 10.


and FEA

(a) 1 Minute

(b) 87 Minutes


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

-5 15 35 55 75 95 115 135 155 175

Tem

pe

ratu

re (

oC

)

Time (min)







(c) 174 Minutes



Test Specimen 10


Mean 1.030 0.976 0.949 0.862

StDev 0.112 0.131 0.191 0.154

CoV 0.109 0.134 0.201 0.178

Overall Mean 0.954

Overall CoV 0.155








Test Specimen 11 consisted of a composite panel formed by sandwiching a layer of

rock fibre insulation between the plasterboards. This specimen has a cavity of 13 mm

that was filled with rock fibre. Boral Firestop plasterboard with 729 kg/m3 in density

was used while the density of the rock fibre was 114 kg/m3. Thermocouples were

located on the specimen as shown in Table 4.1. The specimen was fire tested for about



181 minutes. The behaviour of this specimen was similar to that of Specimen 10. The

test was stopped when the paper on the ambient surface started to burn. Figure 4.39

shows the time-temperature profiles across the plasterboard thickness for Test Specimen

11 and compares them with the results from finite element modelling. Figure 4.40

shows the temperature distributions in the cross-section of Specimen 11.


and FEA

(a) 1 Minute


0 100 200 300 400 500 600 700 800 900

1000 1100 1200 1300 1400 1500

0 20 40 60 80 100 120 140 160 180

Tem

pe

ratu

re (

oC

)

Time (min)







(b) 91 Minutes

(c) 181 Minutes









Test Specimen 11

FS Pb1 Pb2 Amb

Mean 1.105 1.011 1.228 1.046

StDev 0.121 0.150 0.276 0.206

CoV 0.109 0.148 0.225 0.197

Overall Mean 1.097

Overall CoV 0.170



4.17. Summary

This chapter has presented the details of finite element modelling of small scale

plasterboard panels using SAFIR and GID. It also described the details of 11 small scale

fire tests on the thermal performance of plasterboards and insulations and the results.

Finite element analysis (FEA) results were then compared with corresponding

experimental results in this chapter. A good comparison of FEA and experimental

results showed the accuracy of the developed finite element models and associated

thermal properties in simulating the thermal performance of plasterboard panels

including the new composite panel.

Five small scale tests were performed on Boral Firestop plasterboard to study the

thermal performance of multiple boards and their varying thickness. Test result

summary can be seen in Table 4.13. Although the failure temperatures on the ambient

side in each specimen were not the same, all the tests were stopped when the

plasterboard at the ambient side started to burn and all of them showed similar

behaviour.

Table 4.13: Summary of Gypsum Plasterboard Small Scale Test

Test

Specimen

Specimen

Thickness

(mm)

Failure Time

(min)

Ambient

Temperature

(oC)

Exp / FEA results

Overall

Mean

Overall

CoV

1 13 30 266 0.986 0.152

2 16 78 271 1.093 0.12

3 29 171 249 1.044 0.109

4 32 222 249 1.038 0.094

5 48 186 213 0.913 0.124

The time of exposure to the cellulosic fire curve determines the approximate depth up to

which the free and chemically bound water present in the gypsum plasterboard gets

expelled. On average, 1 minute of fire exposure is required to expel water from 1 mm

thickness of plasterboard. Hence in the case of 13 mm thick plasterboard exposed to

standard time-temperature curve from one side, the temperature on the ambient surface

would be maintained at about 100oC up to 13 minutes and in the case of 16 mm

plasterboard it would be maintained for up to 16 minutes. Figure 4.41 shows the



comparison between gypsum plasterboard thickness and its failure time. The difference

in Specimen 5 compared with other specimens is probably due to the unexpected

temperature fluctuation in the furnace.

Figure 4.41: Gypsum Plasterboard Failure Time vs. Thickness

The 7 mm and 13 mm depth temperature profiles of Specimen 3 are seen to display

higher temperatures than the equivalent depth temperature profiles of Specimen 1 at

corresponding times. This is due to the influence of the second plasterboard in

Specimen 3 which blocks the escape of heat and redirects most of it back onto exposed

side plasterboard causing it to heat up faster.

The 8 mm and 16 mm depth temperature profiles of Specimen 4 are seen to display

higher temperature than the equivalent depth profiles of Specimen 2 at corresponding

time due to the heat redirected by the ambient side plasterboard.

The advantage of three layers of plasterboard (Specimen 5) over two layers (Specimens

3 and 4) is observed only during the initial two hours of the test. After two hours the

advantage starts reducing rapidly and by about three hours they are equivalent and

display similar thermal performance (see Figure 4.42). This is because the external

plasterboard of the triple layered specimen fell off. The reason for this is because of the

softened screws bending under the dead weight of the plasterboards making it

equivalent to a double layered specimen.

Specimen 1

Specimen 2

Specimen 3

Specimen 4

Specimen 5

0

50

100

150

200

250

0 10 20 30 40 50 60

Failu

re T

ime

(m

in)

Thickness (mm)



Figure 4.42: Ambient Side Time – Temperature Profiles for All Specimens

Six small scale tests were performed with the insulation materials to study its thermal

performance with varying density. Four were conducted on glass fibre insulation and

two were conducted on rock fibre insulation. Test result summary can be seen in Table

4.14.

Table 4.14: Summary of Glass and Rock Fibre Small Scale Test

Test

Specimen

Specimen

Thickness

(mm)

Failure

Time

(min)

Ambient

Temperature

(oC)

Insulation

Density

(kg/m3)

Exp / FEA Results

Overall

Mean

Overall

CoV

6 64 181 218 21.7 0.959 0.155

7 64 179 221 43.4 1.01 0.134

8 58 200 204 37 1.085 0.145

9 45 185 232 168 0.977 0.174

10 57 174 274 100 0.954 0.155

11 45 181 262 114 1.097 0.17

Test Specimens 6 and 7 have the same thickness but different glass fibre insulation

density. However, their failure time and temperature are almost identical. It was hard to

0

50

100

150

200

250

300

0 50 100 150 200 250

Tem

pe

ratu

re (

oC

)

Time (min)

Specimen 1 Specimen 2 Specimen 3 Specimen 4

Specimen 5 Specimen 6 Specimen 7 Specimen 8

Specimen 9 Specimen 10 Specimen 11



compare Specimens 6 and 7 with Specimens 8 and 9 since they have different insulation

thickness and density. Specimen 8 despite having less thickness and density has a

higher failure time then Specimen 7. This is probably because the slight furnace

temperature drop in Specimen 8 at about 150 minutes (see Figure 4.32). Specimens 9

and 6 have similar failure times despite the fact that Specimen 9 has 8 times more

density than Specimen 6, even though Specimen 9 is 30% thinner than Specimen 6.

This result leads to the conclusion that the differences in the density of glass fibre did

not affect the thermal performance significantly. Regardless of insulation thickness and

density, it was seen that the glass fibre insulation became ineffective at about 700oC

making the composite panels follow similar time-temperature profiles up to the end of

the test.

In the initial stages the glass fibre insulation in Test Specimen 9 was seen to perform

better than other glass fibre insulation in Test Specimens 6 to 8. However, this

advantage was lost by the time the fire side temperature of the insulation reached 700oC.

The temperature on the ambient side of the insulation in all the glass fibre test

specimens (6 to 9) were seen to be in close comparison after the exposed surface on the

insulation crossed 700oC (see Figure 4.42).

Different phenomena were observed with rock fibre insulation. Even though Specimen

11 is thinner than Specimen 10, its failure time is slightly higher. Despite the difference

in its failure time, the thermal performance of Specimens 10 and 11 is seen to be nearly

the same despite the differences in the thickness and density (see Figure 4.42).

The thermal performance difference between glass and rock fibre insulations is

probably because of their ability to resist direct exposed fire. Glass fibre could not

withstand direct fire exposure (see Figure 4.26(b)) while rock fibre remained intact until

the end of the test (see Figure 4.36(b)). Since glass and rock fibre insulations have a low

initial thermal conductivity value, the heat could not pass through the insulation. This

leads to rapid temperature increase on Plasterboard 1 – insulation. These can be

observed in Figures 4.27 to 4.39. This rapid temperature increase causes the

Plasterboard 1 to collapse very quickly. In this condition glass fibre was quickly

consumed by fire and lost its ability to provide any resistance to heat transfer.

Compared to test specimens using multiple layer plasterboards, the composite panels

using insulation materials performed better.

Finite Element Analyses of Load Bearing Wall Panels

Numerical Models to Simulate Thermal Performance of LSF Wall Panels 5-1

Chapter 5


5.1. General

In order to investigate the thermal performance of LSF load bearing wall panels made of

steel studs, gypsum plasterboards, glass fibre and rock fibre insulations, the time-

temperature profiles and distribution from six load bearing wall tests conducted by

Kolarkar (2010) at QUT were considered in this research. These tests included LSF

studs lined with various arrangements of plasterboard and insulation such as single or

dual layers of plasterboard with or without cavity insulation and the new composite

panel with external insulation developed by Kolarkar and Mahendran (2008). Finite

element models of tested LSF wall panels were developed using SAFIR and GID in

order to investigate their thermal performance.

This chapter presents the details of the finite element models of LSF load bearing wall

assemblies. Chapter 4 presented the details of the developed thermal finite element

models of plasterboard assemblies. The thermal properties of materials for finite

element modelling were based on the proposed values in Chapter 3. The predicted time-

temperature profiles from the finite element analyses are compared with corresponding

fire test results in this chapter.

5.2. Test Configuration

Kolarkar (2010) conducted six fire tests on load bearing wall specimens of sizes 2400

mm x 2400 mm to represent a typical wall in a building. The studs were spaced at 600

mm centres and attached to the top and bottom tracks to form the LSF wall panel (see

Figure 5.1(c)). All the studs and tracks were fabricated from galvanised steel sheets

having a nominal base metal thickness of 1.15 mm and a minimum specified yield

strength of 500 MPa. The stud dimensions were 90 x 40 x 15 x 1.15 mm lipped channel

and the track dimensions were 92 x 50 x 1.15 mm unlipped channel (see Figure 5.1).



Table 5.1 gives an overview of the six load bearing wall tests used in Kolarkar‟s (2010)

study.

Table 5.1: Load Bearing Wall Configurations in Kolarkar’s (2010) Fire Tests

No. Configuration Insulation Failure Time

(minutes)

1 None 53

2 None 111

3 Glass Fibre 101

(Cavity Insulation)

4 Rock Fibre

107 (Cavity Insulation)

5 Glass Fibre 181 (Unexpected

furnace failure) (External Insulation)

6 Rock Fibre

136 (External Insulation)

In Kolarkar‟s fire tests, the LSF wall specimens were exposed to heat by a propane

fired gas furnace (see Figure 5.2) from one side only and the time-temperature profiles

at various locations across the thickness of the test specimens were measured by using

metal sheathed thermocouples to help assess their fire performance (see Figure 5.3). To

measure the temperatures at various points on the ambient side an infrared gun was

used. The temperature rise of these thermocouples served as the input to the computer

controlling the furnace heat according to the cellulosic fire curve (Standard time-

temperature curve) given in AS 1530.4 (SA, 2005), which is similar to ISO 834-1

(1999) and ASTM E119 (1995).



(a)

(b)

(a) Stud Section (b) Track Section

(c) LSF Wall Frame

Figure 5.1: LSF Wall Panel



Figure 5.2: Gas Furnace

Fire Exposed Side

Unexposed Side

(a) Thermocouple Locations for 1 x 1 Load Bearing Wall Specimen

Fire Exposed Side

Unexposed Side

(b) Thermocouple Locations for 2 x 2 Load Bearing Wall Specimen with and

without Cavity Insulation

Figure 5.3: Thermocouple Locations for Load Bearing Wall Specimens

Pb 1

Pb 2

Pb 1

Pb 2

Pb 3

Pb4

Cavity

Cavity



Fire Exposed Side

Unexposed Side

(c) Thermocouple Locations for 2 x 2 Load Bearing Wall Specimen with External

Insulation

Figure 5.3: Thermocouple Locations for Load Bearing Wall Specimens

The specimen was also subjected to an axial compression load. For this purpose, a

loading frame (see Figure 5.4 (a)) was specially designed to load the individual studs of

a wall specimen in compression directly from the bottom side. It consisted of two

columns firmly bolted to the ground and a universal beam (UB) connecting the two

columns to form an „H‟ shaped portal frame. A second universal beam was bolted to the

floor. Four jacks (see Figure 5.4 (b)) each of 45 kN capacity were mounted on this beam

at a spacing of 600 mm to load each stud.

(a) Loading frame

Figure 5.4: Loading Frame Arrangement (Kolarkar, 2010)

Pb 1

Insulation

Pb 2

Cavity

Pb 3

Insulation

Pb 4



(b) Loading Arrrangement

Figure 5.4: Loading Frame Arrangement (Kolarkar, 2010)

Figure 5.5: Complete Set-up of Load Bearing Wall Test

5.3. Finite Element Models of Load Bearing Walls

Finite element models of the tested load bearing walls were developed using the same

principles described in Chapter 4 for small scale plasterboard panels. Due to the high

variability of the thermal properties with temperature, of materials within the assembly,

a very small finite element mesh was assigned to better simulate the model. Table 5.2



shows the size of mesh, number of nodes and triangular elements within the model.

Figure 5.6 shows the typical mesh for model with void and insulation cavity.

Table 5.2: Meshing Details

Specimen Size (mm) No. of Nodes No. of Triangle Elements

1 0.01 1936 2594

2 0.01 2900 4514

3 0.01 5637 10722

4 0.01 5637 10722

5 0.01 4329 7360

6 0.01 4329 7360

(a) Load Bearing Wall Specimen with Void Cavity

(b) Load Bearing Wall Specimen with Cavity Insulation

Figure 5.6: Finite Element Mesh of Load Bearing Walls

GID can be used as a post-processor to graphically plot the results of SAFIR analysis.

In the post-processing mode GID is capable of displaying thermal contours, plotting the

temperature history of identified node/element. The locations of nodes are shown in

Figure 5.3. The time-temperature profile from each node was exported to .txt file,

which was later used to compare with experimental results in excel format.

Abbreviations used in naming each node are:

FS : Point at fire surface

Pb1 – Pb2 : Point between plasterboards 1 and 2

Pb1 – Ins : Point between plasterboard 1 and insulation (rock or glass fibre)

Ins – Pb2 : Point between insulation (rock or glass fibre) and plasterboard 2



Pb2 - Cav : Point between plasterboard 2 and cavity (void or glass and rock fibre

insulation)

Cav - Pb3 : Point between cavity (void or glass and rock fibre insulation) and

plasterboard 3

Pb3 – Pb4 : Point between plasterboards 3 and 4

Pb3 – Ins : Point between plasterboard 3 and insulation (rock or glass fibre)

Ins – Pb4 : Point between insulation (rock or glass fibre) and plasterboard 4

Amb : Point at ambient side

5.4. Load Bearing Wall Test Specimen 1

The steel frame shown in Figure 5.1(c) was lined on both sides by single layer of

BORAL Firestop gypsum plasterboard. The plasterboards supplied were 1200 mm x

2400 mm with a thickness of 16 mm and a mass of 729 kg/m3. The sheets were

manufactured to the requirements of Australian Standard AS/NZS 2588 – “Gypsum

Plasterboard” (SA, 1988). K type thermocouple wires were installed to measure the

temperature variations across the wall. They were also attached to the hot flange, web

and cold flange of the stud. A total of 50 thermocouple wires were installed in Test

Specimen 1.

During the test, the Edcar software crashed from 9 to 17 minutes, resulting in the loss of

readings during that period. However, the remaining time-temperature graphs were

plotted accurately. After 3 minutes of starting the furnace, smoke was seen coming out

from the top of the wall specimen. At the end of 11 minutes, smoke and steam were

seen to escape from the top of the wall. By 32 minutes the lateral displacement or

bowing of the wall towards the furnace was prominently noticeable. At the end of 53

minutes, the wall failed to support the applied load and the test was stopped. The cause

of failure of the LSF wall specimen could be attributed to the structural failure of the

frame precipitated by the opening of plasterboard joints and partial collapse of

plasterboard on the fire side.

Figure 5.13 shows the finite element model of Test Specimen 1. The time-temperature

profiles obtained from finite element analyses are presented in Figures 5.7 to 5.12 and

compared with corresponding experimental results.



5.4.1. Plasterboards

Figures 5.7 (a) to (d) show that there is a good agreement between FEA and

experimental results of plasterboards. During the test the exposed plasterboard on the

fire side over Stud 4 had fallen off at the end. Stud 4 was located on the right side of

LSF wall panels. Due to the opening, the plasterboard behind the stud (Pb2) was

severely affected by higher temperature. Due to this occurrence, Figure 5.7 (c) shows

that the time-temperature profile for Pb2 is quite close to time-temperature profile of

Pb1, and thus these time-temperature profiles do not agree well with FEA profiles.

The plasterboard on the ambient side was seen to be in good condition with the paper on

the cavity facing surface burnt only in a few locations, thus maintaining the integrity of

the wall. Insulation failure was also not detected as the temperature on the ambient face

of the unexposed plasterboard was much lower than the insulation failure criteria

(maximum average temperature of 140oC above the ambient or a maximum temperature

of 180oC at any location on the ambient surface) until the end of the test as

recommended by AS 1530.4 (SA, 2005).

(a) Left Side

Figure 5.7: Time - Temperature Profiles of Specimen 1 from FEA & Experiment

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Cavity Experiment - Cavity - Pb2 Experiment - Ambient - Top Experiment - Ambient - Bottom SAFIR - Fire Side SAFIR - Pb1 - Cavity SAFIR - Cavity - Pb2 SAFIR - Ambient



(b) Middle Side

(c) Right Side


0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Cavity Experiment - Cavity - Pb2 Experiment - Ambient SAFIR - Fire Side SAFIR - Pb1 - Cavity SAFIR - Cavity - Pb2 SAFIR - Ambient

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Cavity Experiment - Cavity - Pb2 Experiment - Ambient - Top Experiment - Ambient - Bottom SAFIR - Fire Side SAFIR - Pb1 - Cavity SAFIR - Cavity - Pb2 SAFIR - Ambient



(d) Average


Table 5.3: Comparison of Finite Element Analysis and Experimental Results of

Plasterboards for Test Specimen 1

Left Side

(FEA/EXP) Fire Surface Pb1 - Cavity Cavity - Pb2 Top Amb Bottom Amb

MEAN 0.99 0.84 0.78 1.04 1.07

CoV 0.16 0.12 0.20 0.18 0.13

Overall Mean 0.94

Overall CoV 0.16

Middle Side

(FEA/EXP) Fire Surface Pb1 - Cavity Cavity - Pb2 Ambient

MEAN 0.94 0.86 0.69 1.1

CoV 0.16 0.12 0.20 0.14

Overall Mean 0.90

Overall CoV 0.16

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Cavity Experiment - Cavity - Pb2

Experiment - Ambient SAFIR - Fire Side SAFIR - Pb1 - Cavity

SAFIR - Cavity - Pb2 SAFIR - Ambient





Average

(FEA/EXP) Fire Surface Pb1 - Cavity Cavity - Pb2 Ambient

MEAN 0.97 0.89 0.74 1.06

CoV 0.16 0.14 0.20 0.11

Overall Mean 0.92

Overall CoV 0.15


time-temperature profiles of the plasterboard from FEA with test results. At all the

temperature measuring locations, the correlation between numerical and test results is

reasonably good but is not exact. However, considering software limitations, the

agreement is reasonable. Figures 5.7 (a) to (d) show that the model developed to predict

the time-temperature profiles give good accuracy. Table 5.3 results confirm this with an

average overall mean of 0.92 and an average overall coefficient of variation of 0.15.

5.4.2. Studs

Figures 5.8 to 5.11 show the time-temperature profiles of hot and cold flanges and web

for each stud while Figure 5.12 gives the average time-temperature profiles. They show

that there is a reasonably good agreement between FEA and experimental results of LSF

studs (see Figure 5.12). The reasons for the differences in these results are discussed

next.

The specimen showed no signs of lateral displacement during the initial application of

15 kN compression load. The exposed plasterboard on the fire side over Stud 4 fell off

at the end of the test. Shrinkage of the plasterboard had caused it to detach from the

fasteners, opening the joints and exposing the stud. The joints opened up from 20 mm at

the base to about 35 mm at the top of the stud, indicating the greater severity of the

Right Side

(FEA/EXP) Fire Surface Pb1 - Cavity Cavity - Pb2 Top Amb Bottom Amb

MEAN 0.97 0.97 0.74 0.95 1.16

CoV 0.16 0.19 0.21 0.12 0.13

Overall Mean 0.96

Overall CoV 0.16



plasterboard shrinkage at the top. This was probably due to the higher temperatures in

the chamber at the top due to the upward movement of hot air. SAFIR could not model

the moisture movement, shrinkage and cracking of the wall panels and the predictions

resulted from the FEA are not perfect but are still acceptable. Figure 5.11 shows the

time-temperature profiles for Stud 4. It shows that the top section of Stud 4 has higher

temperatures than the middle and the bottom sections.

The vertical plasterboard joints of the exposed side were on Studs 2 and 4. This resulted

in a much higher temperature of the hot flange at failure when compared to the hot

flanges of Studs 1 and 3. This can be seen in Figures 5.8 to 5.11. SAFIR could not

model the opening of the plasterboard joints caused by the shrinkage of plasterboard at

higher temperatures. This resulted in the uniform temperature across the length of the

model and therefore could not accurately predict the time-temperature profiles for Studs

2 and 4. Figures 5.9 and 5.11 show that the FEA time-temperature profiles are lower

than those from test.

(a) Hot Flange

Figure 5.8: Time - Temperature Profiles of Stud 1 from FEA and Experiment

0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)

SAFIR Experiment - Top Experiment - Middle

Experiment - Bottom Experiment - Average



(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)





(a) Hot Flange

(b) Web


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)


Exoeriment - Bottom Experiment - Average

0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange


(a) Hot Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)





(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

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500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)





(a) Hot Flange

(b) Web


0

100

200

300

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500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)



0

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300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange


Figure 5.12: Average Time - Temperature Profiles of Studs 1 to 4 from FEA and

Experiment

0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Hot Flange Experiment - Web Experiment - Cold Flange

SAFIR - Hot Flange SAFIR - Web SAFIR - Cold Flange




Steel Studs for Test Specimen 1

Stud 1

(FEA/EXP) Hot Flange Web Cold Flange

MEAN 0.94 1.08 0.98

CoV 0.21 0.22 0.23

Overall Mean 1.00

Overall CoV 0.22

Stud 2


MEAN 0.66 0.78 0.78

CoV 0.20 0.20 0.20

Overall Mean 0.74

Overall CoV 0.20

Stud 3


MEAN 0.83 0.82 0.81

CoV 0.19 0.21 0.21

Overall Mean 0.82

Overall CoV 0.20

Stud 4


MEAN 0.75 0.94 0.88

CoV 0.19 0.23 0.23

Overall Mean 0.86

Overall CoV 0.22

AVERAGE


MEAN 0.78 0.89 0.85

CoV 0.19 0.20 0.21

Overall Mean 0.84

Overall CoV 0.20




time-temperature profiles of the studs from FEA with test results. At all the temperature

measuring locations, the correlation between numerical and test results is reasonably

good, but is not exact. However, considering software limitations, the agreement is

reasonable. Figures 5.8 to 5.12 show that the model developed to predict the time-

temperature profiles give good accuracy. Table 5.4 results confirm this with an average

overall mean of 0.84 and an average overall coefficient of variation of 0.20. Figure 5.13

shows the temperature distributions in the cross-section of Test Specimen 1 after 26 and

53 minutes (failure).

(a) 26 Minutes

(b) 53 Minutes (Failure)

Figure 5.13: Temperature Distributions from FEA for Test Specimen 1


For this test specimen, the steel frame was lined on both sides by two layers of BORAL

Firestop gypsum plasterboard. The plasterboards supplied were 1200 mm x 2400 mm

with a thickness of 16 mm and a mass of 729 kg/m3. The sheets were manufactured to

the requirements of Australian Standard AS/NZS 2588 – “Gypsum Plasterboard” (SA,



1988). The first plasterboard layer was installed vertically while the second layer was

installed horizontally. K type thermocouple wires were installed to measure the


and cold flange of the studs. A total of 56 thermocouple wires were installed in Test

Specimen 2.

The initial behaviour of Test Specimen 2 was very much similar to that of Test

Specimen 1. After the smoke and steam escaped by the end of 4 minutes, the specimens

displayed periods of steady burning with little or no smoke or steam. This would happen

after the complete burning of the paper and the complete conversion of water into steam

from the plasterboards. Smoke and steam reappeared with subsequent layers of

plasterboard heating up. At the end of 112 minutes the test was stopped because the

specimen could no longer support the applied load.



experimental results of plasterboards. During the test, the exposed plasterboards (Pb1

and Pb2) though severely calcined were still intact offering protection to the studs. The

screws connecting the plasterboards to the studs were seen to have been pulled through

the plasterboard due to thermal bowing. The ambient side plasterboards (Pb3 and Pb4)

were seen to be in a fairly good condition. The unexposed surface of the specimen

showed no visible signs of wall failure.


time-temperature profiles of the plasterboards from FEA with test results. At all the

temperature measuring locations, the correlation between FEA and test results is


agreement is reasonable. Figure 5.14 shows that the model developed to predict the





(a) Left

(b) Middle


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1000 1100 1200

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Time (min)

Experiment - Fire Side Experiment - Pb1 - Pb2 Experiment - Pb2 - Cavity Experiment - Cavity - Pb3 Experiment - Pb3 - Pb4 Experiment - Ambient - Top Experiment - Ambient - Bottom SAFIR - Fire Side SAFIR - Pb1 - Pb2 SAFIR - Pb2 - Cavity SAFIR - Cavity - Pb3 SAFIR - Pb3 - Pb4 SAFIR - Ambient

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Experiment - Fire Side Experiment - Pb1 - Pb2 Experiment - Pb2 - Cavity

Experiment - Cavity - Pb3 Experiment - Pb3 - Pb4 Experiment - Ambient

SAFIR - Fire Side SAFIR - Pb1 - Pb2 SAFIR - Pb2 - Cavity

SAFIR - Cavity - Pb3 SAFIR - Pb3 - Pb4 SAFIR - Ambient



(c) Right

(d) Average


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1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110

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Time (min)


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Left Side

(FEA/EXP) Fire

Surface

Pb1 -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Pb4

Top

Amb

Bottom

Amb

MEAN 0.99 0.82 0.79 0.73 0.88 0.74 0.77

CoV 0.04 0.15 0.18 0.19 0.17 0.18 0.16

Overall

Mean 0.82

Overall CoV 0.15

Middle Side

(FEA/EXP) Fire Surface Pb1 - Pb2 Pb2 - Cav Cav - Pb3 Pb3 - Pb4 Ambient

MEAN 0.98 0.82 0.78 0.69 0.87 0.76

CoV 0.03 0.14 0.16 0.17 0.19 0.17

Overall Mean 0.82

Overall CoV 0.14

Right Side

(FEA/EXP) Fire

Surface

Pb1 -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Pb4

Top

Amb

Bottom

Amb

MEAN 1.01 0.82 0.83 0.71 0.91 0.81 0.81

CoV 0.03 0.18 0.14 0.17 0.19 0.14 0.14

Overall

Mean 0.84

Overall CoV 0.14

Average

(FEA/EXP) Fire Surface Pb1 - Pb2 Pb2 – Cav Cav - Pb3 Pb3 - Pb4 Ambient

MEAN 0.99 0.82 0.80 0.71 0.89 0.78

CoV 0.03 0.15 0.16 0.17 0.18 0.15

Overall Mean 0.83

Overall CoV 0.14



5.5.2. Studs





next.

On the removal of plasterboards Pb1 and Pb2, it was noticed that the studs had been

laterally displaced at the top end. The friction fit joints provided at the top end of each

stud had failed to prevent the slipping of the studs in the lateral direction. This bending

of the studs about the minor axis near the top portion of the wall caused the screws to

pull out from the plasterboard. The studs also displayed distortional buckling in the top

part of their lengths.

The measured temperatures of the middle part of the studs were higher than the top and

bottom level temperatures. This could be due to the thermal bowing of the wall panels

towards the furnace bringing their central parts closer to the furnace burners, thus

causing the middle part of the wall to heat up faster than the top and bottom parts.

(a) Hot Flange


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(b) Web

(c) Cold Flange


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SAFIR Experiment - Top Experiment - Middle Experiment - Bottom Experiment - Average

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SAFIR Experiment - Top Experiment - Middle Experiment - Bottom Experiment - Average



(a) Hot Flange

(b) Web


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(c) Cold Flange


(a) Hot Flange


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(c) Cold Flange


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(a) Hot Flange

(b) Web


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(c) Cold Flange



Experiment

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Stud 1


MEAN 0.83 0.94 0.85

CoV 0.17 0.15 0.19

Overall Mean 0.88

Overall CoV 0.17

Stud 2


MEAN 0.68 0.74 0.69

CoV 0.23 0.20 0.25

Overall Mean 0.70

Overall CoV 0.23

Stud 3


MEAN 0.76 0.88 0.85

CoV 0.19 0.14 0.17

Overall Mean 0.83

Overall CoV 0.17

Stud 4


MEAN 0.76 0.88 0.85

CoV 0.19 0.14 0.17

Overall Mean 0.83

Overall CoV 0.17

AVERAGE


MEAN 0.73 0.82 0.78

CoV 0.20 0.16 0.20

Overall Mean 0.78

Overall CoV 0.19





measuring locations, the correlation between numerical and test results is reasonably

good but is not exact. However, considering software limitations, the agreement is






(a) 56 Minutes




The construction of Test Specimen 3 was very similar to that of Test Specimen 2. The

only difference was in the use of cavity insulation. The cavity in the wall between the

studs was filled with two layers of 50 mm thick glass fibre (with an original density of

13.88 kg/m3) compressed to 90 mm thickness giving a new density of 15.42 kg/m

3.



Figure 5.21 shows the construction of Test Specimen 3 using glass fibre as cavity

insulation. K type thermocouple wires were installed to measure the temperature

variations across the wall. They were also attached to the hot flange, web and cold

flange of the stud. A total of 56 thermocouple wires were installed in Test Specimen 3.


Specimens 1 and 2. After the smoke and steam escaped by the end of 4 minutes, the

specimens displayed periods of steady burning with little or no smoke or steam. This

would happen after the complete burning of the paper and the complete conversion of

water into steam from the plasterboards. There were periods of thick smoke ensuing

continuously from the specimens for almost 30 to 45 minutes. This probably indicates

the burning of the glass fibre used in the walls. Smoke and steam reappeared with

subsequent layers of plasterboard heating up. Lateral deflections were visible after about

70 minutes. The deflection was initially towards the furnace, but near the end of the test,

a reversal of lateral deflection forcing the wall to bow in the outward direction. At the

end of 101 minutes the test was stopped because the specimen could no longer support

the applied load.

Figure 5.21: Construction of Test Specimen 3



experimental results of plasterboard. During the test, most of the exposed plasterboard

(Pb1) had fallen off whereas plasterboard 2 was still intact though severely damaged.

Upon removal of Test Specimen 2 from the loading frame, both Pb1 and Pb2 were fully

collapsed under their own self weight.



The glass fibre insulation was totally burnt out at the lower right hand portion whereas

the remaining portion of the wall cavity insulation was still intact though it had warped

and shrunk to some extent exposing certain parts of the cavity facing surface of

plasterboard 3. Upon removal of glass fibre insulation, the cavity facing surface of

plasterboard 3 was found to be more or less intact although burnt out at the lower right

hand portion. This was probably due to the completely melt glass fibre insulation at the

lower right corner. Figure 5.22 shows that „Pb1 – Pb2‟ and „Pb2 – Cavity‟ time-

temperature profiles for the right section are lower compared to the left and middle

sections. This is probably due to the glass fibre melting down and allowing the heat to

go through the LSF wall panels. Plasterboard 4 was seen to be in a fairly good

condition. However, it had cracked up horizontally at the middle when the wall failed

by bowing in the outward direction. This crack can cause some heat loss and thus

reduce the accuracy of the finite element model slightly.

(a) Left

Figure 5.22: Time - Temperature Profiles of Test Specimen 3 from FEA & Experiment

0

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800

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1000

1100

1200

0 10 20 30 40 50 60 70 80 90 100

Tem

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)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Pb2

Experiment - Pb2 - Cavity Experiment - Cavity - Pb3

Experiment - Pb3 - Pb4 Experiment - Ambient - Top

Experiment - Ambient - Bottom SAFIR - Fire Side

SAFIR - Pb1 - Pb2 SAFIR - Pb2 - Cavity

SAFIR - Cavity - Pb3 SAFIR - Pb3 - Pb4

SAFIR - Ambient



(b) Middle

(c) Right


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0 10 20 30 40 50 60 70 80 90 100

Tem

pe

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)

Time (min)


Experiment - Cavity - Pb3 Experiment - Pb3 - Pb4 Experiment - Ambient SAFIR - Fire Side SAFIR - Pb1 - Pb2 SAFIR - Pb2 - Cavity


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

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re (

oC

)

Time (min)




(d) Average




Left Side

(FEA/EXP) Fire

Surface

Pb1 -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Pb4

Top

Amb

Bottom

Amb

MEAN 0.95 0.83 0.69 0.84 0.92 0.91 0.88

CoV 0.05 0.14 0.30 0.32 0.27 0.17 0.17

Overall

Mean 0.86

Overall CoV 0.20

Middle Side


MEAN 1.00 0.94 0.77 0.87 0.96 0.88

CoV 0.05 0.17 0.22 0.33 0.30 0.17

Overall Mean 0.90

Overall CoV 0.21

0

100

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1200

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)









Right Side

(FEA/EXP) Fire

Surface

Pb1 -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Pb4

Top

Amb

Bottom

Amb

MEAN 1.04 0.98 0.90 0.92 1.01 0.91 0.96

CoV 0.05 0.16 0.21 0.34 0.27 0.17 0.16

Overall

Mean 0.96

Overall CoV 0.19

Average


MEAN 1.00 0.92 0.79 0.88 0.96 0.91

CoV 0.05 0.15 0.22 0.33 0.28 0.16

Overall Mean 0.91

Overall CoV 0.20




reasonably good but is not exact. However, considering software limitations in

simulating all the variations and complexities in experiments, the agreement is

reasonable. Figure 5.22 shows that the model developed to predict the time-temperature

profiles give good accuracy. Table 5.7 results confirm this with an average overall mean

of 0.91 and an average overall coefficient of variation of 0.20.

5.6.2. Studs


for each stud while Figure 5.27 give the average time-temperature profiles. They show



next.



Studs 1, 2 and 3 failed by local buckling (compressive failure) of the hot flange close to

the mid-height of the wall resulting in the reversal of lateral displacement and causing

the outward movement of the wall. Stud 4 was seen to be relatively undamaged. Local

buckling of Stud 3 was seen to occur between the screws connecting the hot flange to

the plasterboards on the fire side, indicating a good support offered by the plasterboard.

While Stud 3 buckled locally between the screws, Stud 2 buckled locally at the screw

location. The upper and lower tracks supporting the studs were relatively undamaged

and were seen holding the studs firmly in place.

A temperature difference was noticed along the length of the studs. Maximum

temperatures were recorded at mid-height and minimum at the top (see Figures 5.24 to

5.27). After the calcinations of the exposed plasterboards, the temperatures started rising

rapidly in the studs. The presence of cavity insulation shielded the cold flanges from

direct heat introducing a large temperature variation across the depth. These conditions

got worse due to glass fibre partially melting at the right lower section of the wall

panels, i.e. make the temperature variation became even larger.

(a) Hot Flange


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(b) Web

(c) Cold Flange


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(a) Hot Flange

(b) Web


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(c) Cold Flange


(a) Hot Flange


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(b) Web

(c) Cold Flange


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(a) Hot Flange

(b) Web


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(c) Cold Flange



Experiment

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Time (min)







Stud 1


MEAN 0.83 0.98 0.93

CoV 0.17 0.17 0.21

Overall Mean 0.91

Overall CoV 0.18

Stud 2


MEAN 0.79 0.81 0.80

CoV 0.17 0.16 0.17

Overall Mean 0.80

Overall CoV 0.17

Stud 3


MEAN 0.78 0.79 0.78

CoV 0.16 0.14 0.17

Overall Mean 0.78

Overall CoV 0.16

Stud 4


MEAN 0.98 1.09 1.04

CoV 0.23 0.27 0.30

Overall Mean 1.04

Overall CoV 0.26

AVERAGE


MEAN 0.82 0.88 0.86

CoV 0.16 0.16 0.19

Overall Mean 0.85

Overall CoV 0.17





measuring locations, the correlation between FEA and test results is reasonably good

but is not exact. However, considering software limitations, the agreement is






(a) 51 Minutes




The construction of Test Specimen 4 was very similar to that of Test Specimen 3. The

cavity in the wall between the studs was filled with two layers of 25 mm thick rock fibre

(with an original density of 100 kg/m3). Figure 5.29 shows the construction of Test

Specimen 4 using rock fibre as cavity insulation. K type thermocouple wires were



installed to measure the temperature variations across the wall. They were also attached

to the hot flange, web and cold flange of the stud. A total of 56 thermocouple wires

were installed in Test Specimen 4.


Specimens 1 to 3. After the smoke and steam escaped by the end of 4 minutes, the




continuously from the specimens for almost 30 to 45 minutes. This probably indicated

the burning of the rock fibre used in the walls. Smoke and steam reappeared with

subsequent layers of plasterboard heating up. Lateral deflections were visible after about

70 minutes. The deflection was initially towards the furnace, but near the end of the test,

a reversal of lateral deflection forced the wall to bow in the outward direction. At the

end of 107 minutes the test was stopped because the specimen could no longer support

the applied load.




experimental results of plasterboards. During the test, the exposed plasterboard (Pb1

and Pb2) had completely fallen off. The rock fibre insulation was almost fully intact

with only the outer layer of insulation having lost its integrity at several locations. Pb3



and Pb4 have remained in good condition until the end of the test at 107 minutes.

However, Pb4 had cracked up when the wall failed by bowing in the outward direction.

The time-temperature profiles of the plasterboards are shown in Figure 5.30.








(a) Left


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SAFIR - Ambient



(b) Middle

(c) Right


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SAFIR - Ambient



(d) Average




Left Side

(FEA/EXP) Fire

Surface

Pb1 -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Pb4

Top

Amb

Bottom

Amb

MEAN 0.93 0.75 0.86 0.72 0.85 0.86 0.99

CoV 0.05 0.14 0.22 0.35 0.39 0.33 0.24

Overall

Mean 0.85

Overall CoV 0.25

Middle Side


MEAN 1.00 0.85 0.92 0.88 0.80 0.90

CoV 0.04 0.15 0.20 0.39 0.45 0.30

Overall Mean 0.89

Overall CoV 0.25

0

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Time (min)









Right Side

(FEA/EXP) Fire

Surface

Pb1 -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Pb4

Top

Amb

Bottom

Amb

MEAN 1.04 0.91 0.96 0.81 0.92 0.84 1.00

CoV 0.04 0.14 0.17 0.37 0.34 0.33 0.23

Overall

Mean 0.93

Overall CoV 0.23

Average


MEAN 0.99 0.84 0.91 0.80 0.86 0.92

CoV 0.04 0.14 0.19 0.36 0.39 0.28

Overall Mean 0.89

Overall CoV 0.23

5.7.2. Studs





next.

Stud 1 experienced a combination of local compressive failure and torsional buckling of

the hot flange. The local buckling was initiated by screw pull out while the torsional

buckling probably occurred because Pb1 and Pb2 could not provide sufficient lateral

restraint. Pb1 was partially collapsed in that region and Pb2 was severely calcined. This

confirms the sudden rise in temperature of S2-Bottom-HF (see Figure 5.32 (a)). Studs 2

and 3 did not suffer torsional failure but they suffered local compressive failure of the

hot flange at mid-height. Stud 4 was seen to be in good condition. The tracks were seen

to be in good condition and maintained good contact with the studs throughout the fire

test. The temperatures along the stud lengths were seen to be more uniform than Test

Specimen 3.



(a) Hot Flange

(b) Web


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(c) Cold Flange


(a) Hot Flange


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(b) Web

(c) Cold Flange


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(a) Hot Flange

(b) Web


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Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange


(a) Hot Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)





(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)






Experiment



Stud 1


MEAN 0.86 1.18 0.87

CoV 0.19 0.29 0.23

Overall Mean 0.97

Overall CoV 0.24

Stud 2


MEAN 0.79 0.88 0.73

CoV 0.20 0.17 0.18

Overall Mean 0.80

Overall CoV 0.18

0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100

Tem

pe

ratu

re (

oC

)

Time (min)







Stud 3


MEAN 0.79 0.82 0.76

CoV 0.18 0.18 0.18

Overall Mean 0.79

Overall CoV 0.18

Stud 4


MEAN 0.95 1.05 0.88

CoV 0.22 0.29 0.24

Overall Mean 0.96

Overall CoV 0.25

AVERAGE


MEAN 0.83 0.94 0.79

CoV 0.17 0.20 0.18

Overall Mean 0.85

Overall CoV 0.19




but is not exact. However, considering software limitations in simulating all the

variations and complexities in experiments, the agreement is reasonable. Figures 5.31 to

5.35 show that the model developed to predict the time-temperature profiles give good

accuracy. Table 5.10 results confirm this with an average overall mean of 0.85 and an

average overall coefficient of variation of 0.19. Figure 5.36 shows the temperature

distributions in the cross-section of Test Specimen 4 after 51 and 101 minutes (failure).



(b) 51 Minutes




The construction of Test Specimen 5 was similar to that of Test Specimen 3 except for a

single layer of 25 mm thick glass fibre of density 13.88 kg/m3 was used as external

insulation. Figure 5.37 shows the construction of Test Specimen 5 using glass fibre as

external insulation. K type thermocouple wires were installed to measure the


and cold flange of the stud.






continuously from the specimens for almost 30 to 45 minutes. This probably indicates


subsequent layers of plasterboard heating up. Lateral deflections were visible only



towards the end of the test. The deflection was initially towards the furnace, but near the

end of the test, a reversal of lateral deflection forced the wall to bow in the outward

direction. At the end of 181 minutes the test was stopped because the specimen could no

longer support the applied load.




experimental results of plasterboards. During the test, the exposed plasterboard (Pb1

and Pb2) had fallen off exposing Pb3 to direct fire. The external glass fibre insulation

used between Pb1 and Pb2 had completely disappeared leaving only some molten glass

traces along the periphery of the specimen. The furnace malfunctions at 72 minutes

causing the fire curve to drop (see Figure 5.38). Because of this, time-temperature curve

for Pb1-Ins and Ins-Pb2 dropped rapidly. The fire curve was stepped up by manually

operating the furnace from 150 minutes onwards.


time-temperature profiles of the plasterboard from FEA with test results. At all the








(a) Left

(b) Middle


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Insulation Experiment - Insulation - Pb2 Experiment - Pb2 - Cavity Experiment - Cavity - Pb3 Experiment - Pb3 - Insulation Experiment - Insulation - Pb4 Experiment - Ambient - Top Experiment - Ambient - Bottom SAFIR - Fire Side SAFIR - Pb1 - Insulation SAFIR - Insulation - Pb2 SAFIR - Pb2 - Cavity SAFIR - Cavity - Pb3 SAFIR - Pb3 - Insulation SAFIR - Insulation - Pb4 SAFIR - Ambient

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190

Tem

pe

ratu

re (

oC

)

Time (min)

Experiment - Fire Side Experiment - Pb1 - Insulation Experiment - Insulation - Pb2 Experiment - Pb2 - Cavity Experiment - Cavity - Pb3 Experiment - Pb3 - Insulation Experiment - Insulation - Pb4 Experiment - Ambient SAFIR - Fire Side SAFIR - Pb1 - Insulation SAFIR - Insulation - Pb2 SAFIR - Pb2 - Cavity SAFIR - Cavity - Pb3 SAFIR - Pb3 - Insulation SAFIR - Insulation - Pb4 SAFIR - Ambient



(c) Right

(d) Average


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190

Tem

pe

ratu

re (

oC

)

Time (min)


0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190

Tem

pe

ratu

re (

oC

)

Time (min)






Left Side

(FEA /

EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4

Top

Amb

Bottom

Amb

MEAN 0.98 0.77 0.76 0.94 0.74 0.70 0.97 1.02 1.09

CoV 0.06 0.13 0.33 0.22 0.22 0.31 0.15 0.11 0.11

Overall

Mean 0.89

Overall

CoV 0.18

Middle Side

(FEA/EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4 Amb

MEAN 1.02 0.83 0.84 0.89 0.80 0.71 0.99 0.98

CoV 0.04 0.14 0.33 0.25 0.22 0.28 0.15 0.17

Overall

Mean 0.88

Overall CoV 0.20

Right Side

(FEA /

EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4

Top

Amb

Bottom

Amb

MEAN 1.01 0.89 0.99 1.21 0.89 0.82 1.03 1.04 1.09

CoV 0.05 0.14 0.41 0.24 0.18 0.20 0.12 0.11 0.10

Overall

Mean 1.00

Overall

CoV 0.17

Average

(FEA /

EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4 Amb

MEAN 1.00 0.83 0.86 1.01 0.81 0.74 1.00 1.04

CoV 0.03 0.12 0.33 0.22 0.20 0.24 0.13 0.12

Overall

Mean 0.91

Overall CoV 0.17



5.8.2. Studs





next.

During the test, shrinkage occurred on Pb1 and Pb2 resulting in the opening of

plasterboard joints by 10 – 15 mm thus exposing the studs to fire. Local buckling of hot

flange and web occurred near the mid-height of Stud 1. Studs 2 and 3 showed local

compressive failure of the entire cross-section close to the mid-span. Stud 4 was seen to

be undamaged. Torsional and flexural buckling about the minor axis did not occur on all

the studs. This was because of the lateral support provided by the plasterboards.

(a) Hot Flange


0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

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re (

oC

)

Time (min)



0

100

200

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400

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600

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800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





(a) Hot Flange

(b) Web


0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange


(a) Hot Flange


0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

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500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





(a) Hot Flange

(b) Web


0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange



Experiment

0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)







Stud 1


MEAN 1.29 1.04 0.88

CoV 0.17 0.17 0.17

Overall Mean 1.07

Overall CoV 0.17

Stud 2


MEAN 0.90 0.85 0.91

CoV 0.19 0.19 0.18

Overall Mean 0.88

Overall CoV 0.19

Stud 3


MEAN 0.86 0.88 0.91

CoV 0.18 0.16 0.14

Overall Mean 0.88

Overall CoV 0.16

Stud 4


MEAN 1.06 1.12 1.05

CoV 0.19 0.17 0.16

Overall Mean 1.08

Overall CoV 0.17

AVERAGE


MEAN 0.96 0.95 0.92

CoV 0.17 0.15 0.15

Overall Mean 0.94

Overall CoV 0.16












(a) 91 Minutes




The construction of Test Specimen 6 was similar to that of Test Specimen 4 except for a

single layer of 25 mm thick rock fibre of density 100 kg/m3 was used as external

insulation. Figure 5.45 shows the construction of Test Specimen 6 using rock fibre as

external insulation. K type thermocouple wires were installed to measure the




and cold flange of the stud.




happened after the complete burning of the paper and the complete conversion of water

into steam from the plasterboards. There were periods of thick smoke ensuing

continuously from the specimens for almost 30 to 45 minutes. This probably indicated


subsequent layers of plasterboard heating up. Lateral deflections were visible only

towards the end of the test. The deflection was initially towards the furnace, but near the

end of the test, a reversal of lateral deflection forced the wall to bow in the outward

direction. At the end of 112 minutes the test was stopped because the specimen could no

longer support the applied load.




experimental results of plasterboards. During the test the exposed plasterboards (Pb1)

had completely fallen off. Rock fibre insulation that was used between Pb1 and Pb2 had

disintegrated near the lower right corner of the wall. Pb2 also collapsed in this area

exposing Pb3 to direct fire. Rock fibre insulation between Pb1 and Pb2 had undergone

overall shrinking leading to the opening of the joints and exposing Pb2.



(a) Left

(b) Middle


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)




(c) Right

(d) Average


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)


0 100 200 300 400 500 600 700 800 900

1000 1100 1200

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)






Left Side

(FEA /

EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4

Top

Amb

Bottom

Amb

MEAN 0.95 0.78 0.75 0.86 0.74 0.78 0.75 0.95 0.97

CoV 0.06 0.12 0.29 0.23 0.24 0.16 0.09 0.18 0.19

Overall

Mean 0.84

Overall

CoV 0.17

Middle Side

(FEA/EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4 Amb

MEAN 1.01 0.85 0.85 1.03 0.78 0.81 0.70 0.89

CoV 0.04 0.14 0.27 0.21 0.23 0.15 0.13 0.22

Overall

Mean 0.87

Overall CoV 0.17

Right Side

(FEA /

EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4

Top

Amb

Bottom

Amb

MEAN 1.04 0.93 1.02 1.06 0.97 0.85 0.77 0.95 1.00

CoV 0.05 0.16 0.27 0.20 0.17 0.15 0.10 0.17 0.15

Overall

Mean 0.96

Overall

CoV 0.16

Average

(FEA /

EXP) Fire

Surface

Pb1 -

Ins

Ins -

Pb2

Pb2 -

Cav

Cav -

Pb3

Pb3 -

Ins

Ins -

Pb4 Amb

MEAN 1.00 0.85 0.88 0.98 0.83 0.81 0.74 0.95

CoV 0.04 0.13 0.26 0.20 0.20 0.14 0.10 0.18

Overall

Mean 0.88

Overall CoV 0.15










5.9.2. Studs

Torsional and flexural buckling about the minor axis did not occur in all the studs since

they had sufficient lateral support. The tracks were seen to maintain good contact and

connection with the studs. Studs 1, 2 and 3 experienced local buckling of flange and

web elements near the mid-height that led to an outward movement. Stud 4 was seen to

be undamaged.




studs (see Figure 5.51).

(a) Hot Flange

0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)






(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)





(a) Hot Flange

(b) Web


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange


(a) Hot Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)





(b) Web

(c) Cold Flange


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)





(a) Hot Flange

(b) Web


0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)





(c) Cold Flange



Experiment

0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)



0

100

200

300

400

500

600

700

800

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140

Tem

pe

ratu

re (

oC

)

Time (min)







Stud 1


MEAN 0.94 1.06 0.96

CoV 0.20 0.18 0.18

Overall Mean 0.98

Overall CoV 0.19

Stud 2


MEAN 0.84 0.85 0.83

CoV 0.21 0.19 0.17

Overall Mean 0.84

Overall CoV 0.19

Stud 3


MEAN 0.90 0.89 0.89

CoV 0.21 0.19 0.17

Overall Mean 0.84

Overall CoV 0.19

Stud 4

Hot Flange Web Cold Flange

MEAN 1.06 1.12 1.04

CoV 0.23 0.22 0.21

Overall Mean 1.07

Overall CoV 0.22

AVERAGE


MEAN 0.91 0.96 0.91

CoV 0.19 0.18 0.17

Overall Mean 0.93

Overall CoV 0.18












(b) 69 Minutes





5.10. Summary

The chapter has described the thermal performance of six load bearing wall panels. Test

results summary can be seen in Table 5.15. Although the failure temperatures on the

ambient side for each specimen are not the same, all the tests were stopped when they

could no longer maintain the applied load. For load bearing wall tests, the finite element

models using SAFIR could not produce the same accuracy as the small scale tests. This

is mainly because of:

-. Unlike the small scale plasterboard test assemblies which are very simple, the

load bearing wall tests had very complex issues such as moisture movement

across the model cross section, ablation and shrinkage were present in the load

bearing wall panels, which SAFIR finite element model could not simulate.

-. Load bearing tests were stopped following structural failures. Thus the load

bearing wall panels were subjected to cracking, opening of the joint and thermal

bowing before the tests were stopped. On the other hand they were absent or

present at lower levels in small scale plasterboard assemblies. However, the

same thermal properties were used in simulating the behaviour of load bearing

wall panels and small scale plasterboard assemblies. Hence the accuracy in

predicting the time-temperature profiles of the load bearing wall panels was not

as good in comparison with small scale plasterboard assemblies.

-. Thermal bowing reduces the accuracy of finite element modelling since the

specimen deflects towards the furnace due to thermal bowing, thus increasing

the temperature at the beginning of the test. However, near the end of the test,

the specimen deflected in the outward direction, thus reducing the temperature.

These effects could not be simulated in finite element modelling.

-. Gradual spalling of gypsum plasterboard exposed the next layer of LSF wall

panels to direct fire. This could not be simulated by SAFIR finite element

models, thus compromising the accuracy of predicted time-temperature profiles.

-. The opening of plasterboard joints and the cracking due to shrinkage made the

heat transfer faster to the next layer of LSF wall panel behind the joint. This lead

to a non-uniform heat transfer mechanism across the width of the load bearing



wall specimens, ie. higher temperatures in some locations and lower in the

others. This effect could not be simulated by SAFIR finite element models.

-. During the load bearing tests it was found that some sections of the plasterboard

or insulation had fallen off. This resulted in uneven heating and thus

compromised the accuracy of predicted time-temperature profiles by SAFIR

finite element models.

-. Water evaporating from the exposed plasterboard moves to the ambient side.

During this moisture movement the layer behind the exposed plasterboard gets

additional humidity, thus changing its thermal property. This could not be

modelled by SAFIR, thus compromising the accuracy of predicted time-

temperature profiles. Thus the moisture movement becomes important in

modelling the tests with void cavity since it plays a major role in heat transfer.

This can be seen in the steel stud cavity figures of Tests 1, 2, 5 and 6. SAFIR

finite element models predicted the steel stud temperatures less accurately for

Tests 3 and 4.

Despite these limitations, time-temperature profiles predicted by SAFIR finite element

models achieved a reasonable agreement with experimental results of load bearing tests.


Load Bearing Wall Panels

Plasterboards

Specimen

1

Specimen

2

Specimen

3

Specimen

4

Specimen

5

Specimen

6

Overall

Mean 0.97 0.84 0.91 0.89 0.90 0.87

Overall CoV 0.18 0.17 0.20 0.23 0.19 0.16

Steel Studs

Specimen

1

Specimen

2

Specimen

3

Specimen

4

Specimen

5

Specimen

6

Overall

Mean 0.88 0.84 0.85 0.85 0.95 0.95

Overall CoV 0.21 0.19 0.17 0.19 0.20 0.21



Following is a list of the main findings from load bearing wall Test Specimens 1 to 6.

-. Compared with cavity insulated wall specimens (Specimens 3 and 4), externally

insulated wall specimens (Specimens 5 and 6) give more protection to the steel

studs and thus have a higher failure time (see Figures 5.53 and 5.54). Figure

5.53(a) shows that higher hot flange temperatures were reached at earlier times

for cavity insulated wall specimens in comparison to externally insulated

specimens. This is probably because the main function of insulation materials is

to eliminate the heat transfer across the wall cavity by radiation and convection,

which are essentially the faster modes of heat transfer in comparison to

conduction. No cavity insulation can reduce the heat transfer towards the cold

flange by conduction along the cross-section of the stud. Thus the cold flange

receives heat from the hot flange by conduction along the web, which would be

the fastest mode of heat transfer in the case of cavity insulated materials.

Because of the very low conductivity of the cavity insulating material as

compared to steel, most of the heat gets directed and channelled along and

across the steel studs which act as the heat sink thus raising their body

temperature much faster than in the case of non-cavity insulated specimens (see

Figure 5.53).

-. Glass fibre cavity insulation lead to a higher hot flange stud temperature

compared with rock fibre cavity insulation. This is probably because, in the

cavity insulated specimens, the insulation is on the ambient side of the hot

flange and thus incapable of offering any protection. Since glass fibre blocks

heat better than rock fibre, the heat cannot pass through from hot flange to the

insulation, thus accelerating the temperature increase in the hot flange.

-. In the case of externally insulated specimens, it is seen that the temperature

profiles of the studs are well separated implying the effect of insulation on the

stud temperatures. Rock fibre insulation gives better protection to the studs

compared with glass fibre insulation.

-. The ambient side temperatures of all the wall specimens were observed to be

below the insulation failure temperatures (maximum average temperature of

140oC above the ambient or a maximum temperature of 180

oC at any location on



the ambient surface as recommended by AS 1530.4 (SA, 2005)). The specimen

failure was always by the structural failure of the studs and never by insulation

or integrity failure. In the cavity insulated specimens, the fire exposed

plasterboard collapsed partially prior to stud failure thus hastening the collapse

of the wall by exposing the steel frame to direct furnace heat. The failure time of

externally insulated specimens was found to be the maximum.

-. Although Test Specimen 5 had problems during the test due to furnace failure,

the finite element model was able to give a good prediction of the time-

temperature profiles for the gypsum plasterboards and steel studs.

(a) Hot Flange

Figure 5.53: Time - Temperature Profiles from FEA for Steel Studs

0

50

100

150

200

250

300

350

400

450

500

0 50 100 150 200

Tem

pe

ratu

re (

oC

)

Time (min)





(b) Web

(c) Cold Flange

Figure 5.53: Time - Temperature Profiles from FEA for Steel Studs

0

50

100

150

200

250

300

350

400

450

500

0 50 100 150 200

Tem

pe

ratu

re (

oC

)

Time (min)



0

50

100

150

200

250

300

350

400

450

500

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





Figure 5.54: FEA Time - Temperature Profiles for Ambient Side Gypsum

Plasterboards

5.11. Improving Composite LSF Wall Panel

To further improve the composite LSF wall panel, SAFIR models can be used as a

reference on how the new cross section will react if exposed to standard fire test. Based

on the discussions in the last section and Figure 5.53, rock fibre insulation was found to

provide better protection to steel studs than glass fibre insulation. Since the specimen

failure was always by the structural failure of the studs and never by the insulation or

integrity failure, rock fibre insulation is a better choice for improving the composite

wall panel. The other reason is that rock fibre insulation is more able to resist

disintegration by direct fire exposure compared with glass fibre insulation. External

insulation is a better choice then cavity insulation since it gives more protection to the

steel studs.

0

10

20

30

40

50

60

70

80

90

100

0 20 40 60 80 100 120 140 160 180 200

Tem

pe

ratu

re (

oC

)

Time (min)





Figure 5.55: New Composite Panel – FEA Experiment 1 (FEA1)

where:

= Steel

= Gypsum Plasterboard

= Rock Fibre Insulation

Figure 5.55 shows a new composite wall panel used in finite element modelling to

investigate if a new section will improve the fire resistance rating of externally insulated

wall panels. The thermal properties were the same as those used in Specimen 6 (rock

fibre external insulation). The overall thickness of the wall panel is also the same as

Test Specimen 6. The new composite wall panel cross section consists of three 13 mm

plasterboards with two 9 mm rock fibre insulation giving a total thickness of 57 mm on

each side.

Finite element model of the new composite panel was developed using the same

principles described in Chapters 4 and 5. It was then subjected to standard fire

conditions. Due to the high variability of the thermal properties with temperature of the

materials used in the assembly, a very small finite element mesh was assigned to better

simulate the model (0.01 in mesh size).

Table 5.16 shows the failure temperatures of steel studs and ambient side plasterboards

based on finite element modelling while Tables 5.17 shows the failure temperatures

based on Kolarkar‟s (2010) load bearing wall tests. The FEA results will be used as a



benchmark to measure if the new cross section could perform better than the current

composite wall panel.

All the specimens started to fail when the temperature of the hot flange reaches around

423oC (average hot flange temperature of all specimens). At that temperature, the steel

starts to lose its strength and buckle under axial loading, leading to structural failure of

the whole wall panel specimen. Since Test Specimen 5 unexpectedly experienced

furnace failure, it could not be used as a benchmark for further improvements, and

therefore Test Specimen 6 was used instead.

Table 5.16: Failure Temperature of Steel Studs and Ambient Side Plasterboards

from FEA

Specimen Ambient

(oC)

Hot Flange

(oC)

Web

(oC)

Cold Flange

(oC)

Time

(min)

1 93 447 407 376 52

2 56 385 357 335 111

3 56 416 303 196 101

4 55 479 332 205 107

5 55 422 398 375 183

6 42 393 363 342 137

Table 5.17: Failure Temperature of Steel Studs and Ambient Side Plasterboards

from Experiments (Kolarkar, 2010)

Specimen Ambient

(oC)

Hot Flange

(oC)

Web

(oC)

Cold Flange

(oC)

Time

(min)

1 72 525 416 375 52

2 69 465 398 375 111

3 54 511 352 229 101

4 50 475 264 197 107

5 53 496 433 410 183

6 57 466 390 357 137



Figure 5.56: Comparison between FEA1 and Test Specimen 6 (SAFIR)

Table 5.18: Failure Temperatures of FEA1

Hot Flange (oC) Web (

oC) Cold Flange (

oC) Ambient (

oC) Time (min)

Specimen 6 393 363 342 42 137

FEA1 392 369 350 47 186


oC) Cold Flange (

oC) Ambient (

oC) Time (min)

Specimen 6 393 363 342 42 137

FEA1 385 361 342 47 184

Hot Flange (oC) Time (min)

Average 423 -

FEA1 422 194

From Table 5.18 it can be seen that the new composite wall panel gives better

protection to the steel studs. If the hot flange temperature governs the failure criteria,

there is an increase of 36% in the fire performance of the new wall panel in terms of

failure time. If the cold flange temperature governs the failure criteria, there is an

0

200

400

600

800

1000

1200

1400

0 50 100 150 200 250 300 350

Tem

pe

ratu

re (

oC

)

Time (min)

SAFIR - FS SAFIR - Amb SAFIR - HF SAFIR - Web SAFIR - CF

FEA1 - FS FEA1 - Amb FEA1 - HF FEA1 - Web FEA1 - CF



increase of 34% in the fire performance of the new wall panel in terms of failure time.

Either way, the new LSF Wall Panel gives considerable fire rating improvement.

From Figure 5.56, it can be seen that the temperatures of hot flange, web, and cold

flange of FEA1 are gradually merging. This is probably because the heat could not pass

through the ambient side. The improvement in the thermal performance of the new LSF

panel was due to the use of double layers of rock fibre insulation. When the fire exposed

plasterboard (Pb1) completely falls off, the first layer of rock fibre will be subjected to

direct fire exposure. Since rock fibre has very good resistance against disintegration it

will protect the second layer of plasterboard (Pb2) from direct fire exposure. Thus

spalling off of gypsum plasterboard will be prevented at least until the rock fibre

completely falls off. This will be repeated by the second layer of rock fibre and the third

layer of gypsum plasterboard. Figure 5.57 shows the time-temperature profiles of the

new LSF composite wall panel.

To further improve the wall panel, thin steel sheets were added between the studs and

the plasterboard. This will allow some heat to get through the ambient side quicker

(conduction heat transfer) and further reduce the steel temperature inside the cavity,

leading to further increase in the failure time. The addition of steel sheet will also

increase the structural strength of walls. It will help reduce the occurrence of local

buckling of hot and cold flanges and add lateral strength. Because of the increase in the

structural strength, the failure temperature of the hot flange will more likely exceed

423oC. Further research is needed to investigate how the wall panel will collapse. For

this research the failure time will be investigated until the hot and cold flanges reach

393oC and 342

oC, respectively (same as Specimen 6).



Figure 5.57: Time - Temperature Profiles of FEA1

Figure 5.58: New Composite Panel - FEA Experiment 2 (FEA2)

where:

= Steel

= Gypsum Plasterboard

= Rock Fibre Insulation

0

200

400

600

800

1000

1200

0 50 100 150 200 250 300

Tem

pe

ratu

re (

oC

)

Time (min)

FS Pb1 - Ins Ins - Pb2 Pb2 - Ins Ins - Pb3 Pb3 - Cav

1 mm Steel Sheet



Figure 5.58 shows another composite wall panel used in finite element modelling to

investigate if it will further improve the fire rating of externally insulated wall panels.

The thermal properties used were the same as those used in Specimen 6 (rock fibre

external insulation). The overall thickness of the wall panel is also the same as Test

Specimen 6. The FEA2 wall panel cross section consists of a layer of 1 mm thick steel

sheet, two 13 mm plasterboards and one 12 mm plasterboards (near cavity) with two 9

mm rock fibre insulation giving a total thickness of 57 mm on each side. Finite element

model was developed using the same principles described in Chapters 4 and 5. It was

subjected to standard fire conditions. Due to the high variability of the thermal

properties with temperature within the assembly, a very small finite element mesh was

assigned to better simulate the model (0.01 in mesh size).

Table 5.19: Failure Temperatures of FEA1 & FEA2


oC) Cold Flange (

oC) Ambient (

oC) Time (min)

Specimen 6 393 363 342 42 137

FEA1 392 369 350 47 186

FEA2 393 365 343 48 194


oC) Cold Flange (

oC) Ambient (

oC) Time (min)

Specimen 6 393 363 342 42 137

FEA1 385 361 342 47 184

FEA2 393 365 343 48 194

Hot Flange (oC) Time (min)

Average 423 -

FEA1 422 194

FEA2 422 204



Figure 5.59: Time - Temperature Profiles of Steel Studs

From Table 5.18 and Figure 5.59 it can be concluded that adding steel sheets between

studs and plasterboards make the heat travel faster to the ambient side. It will make the

temperature inside the cavity lower and increase the temperature on the ambient side. It

will also makes the system to have a longer failure time and improve the fire rating of

the wall panel.

0

100

200

300

400

500

600

700

0 50 100 150 200 250 300 350

Tem

pe

ratu

re (

oC

)

Time (min)

SAFIR - Amb SAFIR - HF SAFIR - Web SAFIR - CF

FEA1 - Amb FEA1 - HF FEA1 - Web FEA1 - CF

FEA2 - Amb FEA2 - HF FEA2 - Web FEA2 - CF

Conclusions and Recommendations


Chapter 6


6.1. Conclusions

This research was conducted to develop suitable finite element models of load bearing

LSF wall and small scale plasterboard panels consisting of gypsum plasterboard, steel,

glass and rock fibre insulation materials and investigate their thermal behaviour under

standard fire conditions. These panels included both conventional LSF wall systems

with and without cavity insulation and the new externally insulated composite panel

system. For this purpose, SAFIR finite element program and GID pre and post

processors were used to develop suitable numerical models and investigate the thermal

behaviour of eleven small scale plasterboard models and six load bearing wall

assemblies.

Tests were also conducted to measure the thermal properties of gypsum plasterboard,

glass and rock fibre insulations. A review of the thermal properties as reported by

various researchers in this field of study was also included as a reference to develop

idealised thermal properties to be used in finite element modelling. The small scale

plasterboard panel test results were used to develop the idealised thermal properties of

each material used in this study, since complex issues such as moisture movement

across the model cross section, ablation and shrinkage are less important in the small

scale plasterboard panels. The fire test results of these small scale plasterboard panels

were used first in the comparison with finite element analysis results to confirm the

accuracy of idealised thermal properties. The same thermal properties were then used

with the developed numerical models of LSF load bearing wall assemblies. A detailed

comparison of time-temperature curves predicted by the developed finite element

models and available fire test results (Kolarkar, 2010) was undertaken for both small

scale plasterboard models and load bearing wall assemblies to determine the accuracy of

finite element models.

It was found that the developed finite element models with the proposed thermal

properties of gypsum plasterboard, glass and rock fibre insulations were able to

accurately predict the thermal response of small scale plasterboard models subject to



standard fire conditions. However, in the case of load bearing wall assemblies, the time-

temperature profiles predicted by the developed finite element models did not match as

well as in the case of small scale plasterboard models. However, considering the more

complicated behaviour of larger wall assemblies, the agreement between numerical and

experimental results is considered reasonable. The reasons for the differences between

finite element analysis and fire test results are:

-. Non-uniform heating of furnace was present in the load bearing wall tests, which

resulted in variations of time-temperature profiles. In comparison, the small

scale tests only used one burner, thus resulting in uniform heating across the

entire test specimen.

-. The load bearing wall test was stopped following a structural failure of one or

more steel studs whereas the small scale test was stopped when the ambient side

plasterboard commenced burning. Hence there were more complicated and

larger deformations in the load bearing wall assemblies during the test, which

affected the thermal performance, but were not captured by the developed two-

dimensional finite element models.

-. Moisture movement in the void cavity of load bearing wall panel could not be

modelled accurately by SAFIR finite element models.

-. Load bearing wall panels consisted of several plasterboards with joints between

them. These joints opened up during the standard fire tests, which affected the

measured time-temperature profiles. Effects of such joints could not be included

using the developed finite element models. Small scale plasterboard model tests

consisted of only single plasterboard without any joints.

The thermal properties of gypsum plasterboard, glass and rock fibre insulations were

calibrated to accommodate the effects of spalling of gypsum plasterboard and thermal

bowing since these phenomenon occurred in both cases. Despite the differences

between small scale and load bearing wall tests, this research used the same thermal

properties for small scale plasterboard models and load bearing wall assemblies. Further

research is needed to advance the capabilities of finite element software to be able to

simulate the effects of moisture movement, ablation, cracking and shrinkage that occur

in LSF wall panels during fires.



A comparison of finite element model predictions versus fire test results in our research

and past research conducted on LSF wall panels by others show that the agreement has

improved noticeably in our research. This is due to the following reasons: the use of

more appropriate thermal properties based on testing, literature review and calibration,

and improved accuracy in the temperature measurements during fire tests.

Using the developed finite element model of load bearing LSF wall panels, this thesis

has also investigated the thermal performance of various LSF wall configurations and

showed that the composite panel system proposed by Kolarkar and Mahendran (2008)

performs better in fires than other panel systems such as those with cavity insulation.

Using these models, it also proposed a new composite wall panel system consisting of

three layers of plasterboard and two layers of rock fibre with thin steel sheets in the void

cavity. The new composite wall panel system is capable of increasing the fire resistance

rating by 30% in comparison with the current composite wall panel.

In summary, the developed numerical models using SAFIR and GID can be used to

simulate the thermal behaviour of LSF wall systems including that with the new

composite panel under fire conditions with an acceptable accuracy. It is particularly

useful in comparing the thermal performance of different wall panel systems without

time consuming and expensive full scale fire tests. It can also be easily used to further

improve the cross section of the wall panels without doing expensive tests.

6.2. Recommendations for Future Research

Further improvements are needed in the area of finite element modelling in order to

obtain more accurate results. The capabilities of available finite element programs such

as SAFIR should be further improved to include the effects of:

-. Ablation of plasterboard

-. Shrinkage of plasterboard

-. Opening of joints and cracking of plasterboard

-. Removal of materials (glass fibre melting, glass fibre collapse)

-. Moisture movement across the cross section

-. Thermal bowing

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