c 2017 emerald publishing limited notice changes ...abstract: load-bearing light gauge steel frame...
TRANSCRIPT
This is the author’s version of a work that was submitted/accepted for pub-lication in the following source:
Kesawan, Sivakumar & Mahendran, Mahen(2017)Fire performance of LSF walls made of hollow flange channel studs.Journal of Structural Fire Engineering, 8(2).
This file was downloaded from: https://eprints.qut.edu.au/105430/
c© 2017 Emerald Publishing Limited
Notice: Changes introduced as a result of publishing processes such ascopy-editing and formatting may not be reflected in this document. For adefinitive version of this work, please refer to the published source:
https://doi.org/10.1108/JSFE-03-2017-0027
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Fire Performance of LSF Walls made of Hollow Flange Channel Studs
Sivakumar Kesawan and Mahen Mahendran
School of Civil Engineering and Built Environment, QUT, Brisbane, Australia
Abstract: Load-bearing Light gauge Steel Frame (LSF) walls are commonly made of cold-
formed lipped channel section (LCS) studs, and lined with gypsum plasterboard layers.
Recent research on LSF wall systems using full scale fire tests has shown that the use of
welded Hollow Flange Channel (HFC) studs exhibited superior fire performance than those
made of LCS. However, comprehensive fire performance data is not available for HFC stud
walls as only five fire tests were conducted. To advance the use of HFC studs, a detailed
parametric study was performed using validated finite element models to investigate the
structural fire performance of HFC studs subject to non-uniform temperature distributions.
The effects of stud sizes, stud profiles, elevated temperature mechanical property reduction
factors of steels, wall configuration, plasterboard to stud connectivity and realistic design fire
curves were evaluated. The load ratio versus Fire Resistance Rating (FRR) curves were
produced that enabled an easier comparison of the effects of different parameters. This paper
presents the details of this parametric study and the results. It also evaluates the applicability
of the critical temperature methods to predict the FRR of LSF walls made of HFC sections.
Keywords: LSF walls, Hollow flange channel sections, Fire resistance rating, Structural fire
behaviour, Finite element modelling, Parametric study,
Corresponding author’s email address: [email protected]
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1. Introduction
Load-bearing Light gauge Steel Frame (LSF) walls are fabricated using cold-formed steel
frames, and are lined with gypsum plasterboard layers on both sides. When exposed to fire
conditions, these walls must withstand the applied load for a certain period without a
structural failure to provide the required fire resistance rating (FRR). For the fire design of
LSF walls, engineers rely on the FRR (minutes) given by the plasterboard manufacturers,
which are limited to the parameters used in their full scale fire tests. This inhibits
advancements in LSF walls as new stud sections or wall configurations cannot be used
without further, more expensive and time consuming, full scale fire tests. This is due to the
lack of sound knowledge on the effects of important parameters such as stud sizes and
profiles on the fire performance of LSF walls. This has also hindered the process of
developing a performance based approach for the fire design of LSF walls.
Conventionally the LSF walls are made of Lipped Channel Section (LCS) studs (Fig.1(a)).
Recent research at QUT recommended using a welded Hollow Flange Channel (HFC) (Fig.
1(b)) as studs in LSF walls [1]. These HFC sections have higher local and distortional
buckling capacities due to the absence of free edges and the presence of two torsionally rigid
hollow flanges. The structural efficiencies of HFC sections at ambient temperature have been
demonstrated in recent research studies [2-4]. Further, if LSF walls are made of HFC section
studs, the connectivity between plasterboards and steel studs is improved as the connecting
screws can penetrate through both inner and outer flanges (Fig. 1b). Kesawan and Mahendran
[1] conducted five full scale fire tests of LSF walls with studs made of a welded HFC section
known as LiteSteel Beam (LSB) (Fig. 2). It was found that these walls provided more than
50% improvement to FRR in comparison to walls made of LCS. Similarly Jatheeshan and
Mahendran [5] demonstrated the superior fire performance of LSF floors made of the same
welded HFC sections as joists. However, a comprehensive database on the fire performance
of LSF walls made of HFC studs could not be developed through these fire tests as they are
expensive and time consuming. Therefore Kesawan and Mahendran [6] developed suitable
finite element models to predict the thermal and structural performances of LSF walls, and
validated them by using their full scale fire test results. The developed models of HFC section
studs subjected to non-uniform temperature distributions in LSF walls predicted the time-
temperature profiles and structural capacities with good accuracy. These models were then
used in a detailed parametric study to develop a good understanding of the effects of many
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relevant parameters such as stud sizes including thickness and depth, stud profiles and stud to
plasterboard connectivity on the performance of LSF walls exposed to standard fire
conditions [7]. Effects of different wall configurations, elevated temperature mechanical
property reduction factors and realistic design fire curves were also evaluated. Non-
dimensional load ratio versus FRR (failure times) curves were produced for all the LSF walls
considered in this study. This paper presents the details of this parametric study on the fire
performance of LSF walls made of HFC studs and the results. It also includes a brief
summary of the developed models to predict the structural fire behaviour of HFC section
studs. The applicability of the critical temperature method was also investigated for LSF
walls made of HFC studs, and the results are presented in this paper.
2. Finite Element Model Development
Previous researchers have successfully used the finite element analysis program ABAQUS to
predict the structural capacity of thin-walled members including the LCS and HFC sections at
ambient temperature [2, 8]. Previous researchers [9-13] used ABAQUS to predict the
structural capacity of LCS studs subject to non-uniform temperature distributions in fires.
The structural behaviour of HFC sections is different to those of LCSs due to the absence of
free edges and the presence of rigid hollow flanges. Kesawan and Mahendran [6] developed
suitable finite element models to predict the structural capacity of the failed HFC section
studs in their fire tests of load bearing LSF walls. They used the measured stud dimensions
and time-temperature profiles from the fire tests together with the elevated temperature
mechanical property reduction factor models proposed by Kesawan [14]. Table 1 presents the
details of the five fire tests and the results from tests and finite element analyses (FEA). As
evident from Table 1, their FEA predictions agreed well with the fire test results. The
techniques adapted in Kesawan and Mahendran’s [6] models were used to develop the
models used in this parametric study, which are discussed next.
2.1. Hollow Flange Section Stud Model
LSB is a welded HFC section (LiteSteel Beam) used by Kesawan and Mahendran [1] in their
fire tests. Built-up sections shown in Fig. 1(c) can also be manufactured cost effectively using
the available G250 and G500 cold-formed steel sheets in Australia. Zhao and Mahendran’s
[15] research showed that if their recommended suitable screw diameters and spacings are
used to make such screw or rivet fastened HFC sections, the behaviour and strength of such
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built-up sections will be very similar to the fully welded HFC sections. Hence in this section
finite element models of welded HFC sections made of G500 steel (minimum yield strength
500 MPa) were developed.
2.2. Element Types and Boundary Conditions
Shell elements of S4R type were used to model the wall stud while rigid plates made of
R3D4 elements were attached to each end of the stud (Fig. 3). The mesh size used was 4 mm
x 4 mm. Pinned boundary conditions were defined at the geometrical centroids on both ends.
Fire tests [1] showed that plasterboards provide lateral restraints to the wall studs, resulting in
the elimination of lateral torsional and minor axis buckling failures. In the developed models,
the lateral restraints available to the studs were considered until failure by resisting the stud
movement along the minor axis/Z direction at both the inner and outer flanges (Fig. 3) as the
screws connected the plasterboards to the studs by penetrating through both flanges. The wall
height/stud length was maintained as 2.4 m in this parametric study.
2.3. Temperature Development in LSF Walls
Kesawan and Mahendran [16] developed thermal finite element models to predict the time-
temperature profiles of LSF walls under standard fire conditions [7] using SAFIR, and
validated them using full scale fire test results. Effects of plasterboard joints were not
included in these models. The validated thermal FE models were then used to determine the
time-temperature profiles of LSF walls made of HFC studs with varying sizes and section
profiles and wall configurations. These time-temperature profiles were used as input to
ABAQUS analyses. The temperature distribution along the flange widths was taken as
uniform while it was taken as linear across the web depth and lip as shown in Fig. 4 based on
Kesawan and Mahendran’s [6] findings.
2.4. Wall Configurations
Different types of LSF wall configurations are used in buildings. In this study, three wall
configurations; Uninsulated LSF walls lined with dual plasterboard layers - Configuration A,
Uninsulated LSF wall lined with single plasterboard layer - Configuration B and Cavity
insulated LSF walls lined with dual plasterboard layers - Configuration C, as shown in Fig. 5
were considered as they are commonly used in Australia.
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2.5. Elevated Temperature Mechanical and Thermal Properties of Steel
The use of suitable mechanical property reduction factors is imperative for accurate FRR
predictions. Elevated temperature yield strength and elastic modulus reduction factors given
by previous researchers differ largely among them due to the differences in the
manufacturing process, chemical compositions, thickness and grade [14]. In this study three
different steel types, Types A and B – High strength (G500) and Low strength (G250) steels
with Dolamune Kankanagme and Mahendran’s [17] mechanical property reduction factors
and Type C - G250 steel with Eurocode 3 Part 1.2 mechanical property reduction factors
(Tables 2 and 3), were considered. Dolamune Kankanamge and Mahendran’s [17] models
were selected as they were based on the typically used cold-formed steel sheets in Australia.
The nominal ambient temperature yield strengths of G500 and G250 steels were taken as 500
and 250 MPa, respectively, while the elastic modulus was taken as 200,000 MPa for both
steel types. Further, elevated temperature stress-strain relationships of cold-formed steels
were obtained using Dolamune Kankanamge and Mahendran’s [17] predictive equations
based on Ramberg and Osgood’s [18] model. The elevated temperature relative thermal
elongation (Δl/l) values used were from Eurocode 3 Part 1.2 [19].
2.6. Initial Geometric Imperfections and Residual Stresses
The member strengths are sensitive to the imperfections in the shape of their eigen modes
[20]. In this study the initial geometric imperfection of HFC stud was inserted in the shape of
its critical eigen mode. Based on AS 4100 [21], an amplitude of b/150 was used for the local
geometric imperfection, where b is the web height, while an amplitude of l/1000 was used for
the global geometric imperfection, where l is the column length (Figs. 6(a) and (b)). Residual
stresses were not included as they have negligible influence on the structural capacity of HFC
studs even at ambient temperature, and further they diminish at elevated temperatures [6].
2.7. Analysis Method and Comparison of Results
In the previous studies [10, 11, 22] the steady state analysis method was successfully used in
the numerical simulations of LCS studs. Kesawan and Mahendran’s [6] simulation results
using the transient and steady state methods agreed well with the fire test results (Table 1).
Hence in this study, the steady state analysis method was used as it needs less memory and
time. In this method, the stud temperatures were increased first to the target temperatures, and
the load was then increased until failure.
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The failure load obtained from FEA was used to determine the load ratio, which is the ratio of
the structural capacities of HFC stud under fire and ambient conditions. The load ratio versus
FRR curve was then produced for each LSF wall considered. Following sections present the
FEA results as a function of each parameter investigated and associated discussions on its
effects on the fire performance of LSF walls. The non-dimensional nature of these curves
enabled easier comparisons between the FRR of different LSF walls.
3. Effect of Stud Depth on the Fire Performance of LSF Walls
The effect of stud depth on the fire performance of LSF walls was investigated by
considering three HFC sections with varying web depth; 60x45x15x1.6, 90x45x15x1.6 and
150x45x15x1.6 mm (Fig. 7) made of G500 steel (Type A). Their finite element models were
developed using the procedures given in Section 2, and the results are evaluated and
discussed next.
Fig. 8 shows the time-temperature profiles of the three HFC studs while Figs. 9 and 10 show
the load ratio versus failure time (FRR) and critical outer hot flange temperature curves. The
load ratio does not drop until 20 minutes for Wall Configuration A (Fig. 9(a)) as the stud
temperatures were closer to the ambient temperature (Fig. 8). Thereafter increasing hot and
cold flange temperatures led to reduction in the fire performance of LSF walls as the
mechanical properties of steel deteriorated at elevated temperatures. There were significant
differences between the FRR of LSF walls made of studs with varying section depths.
Fig.9(a) results show that at a load ratio of 0.6, LSF wall with Configuration A made of
60x45x15x1.6 mm stud had a FRR of 100 minutes while those made of 90x45x15x1.6 and
150x45x15x1.6 mm studs had FRRs of 120 and 140 minutes, respectively. This difference in
FRR was observed although there was hardly any effect on the temperature development of
steel stud due to varying stud depth as seen in Fig. 8. Further, as seen in Fig. 10(a), the load
ratio versus critical outer hot flange temperature (at failure) curves are dissimilar for LSF
walls made of studs with varying depths, i.e., the critical outer hot flange temperatures were
300, 375 and 450oC at a load ratio of 0.6 for 60x45x15x1.6, 90x45x15x1.6 and
150x45x15x1.6 mm studs, respectively.
The differences in FRR and critical hot flange temperature are due to two main reasons; the
influence of thermal bowing deformation and the effect of failure mode. The temperature
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difference between the hot and cold flanges induces thermal bowing deformation, which is
inversely proportional to the section depth. Hence studs with smaller depth are subject to
higher bending actions and thus lead to reduced FRR. The failure mode could also affect the
fire performance of LSF walls. 150x45x15x1.6 mm stud was subjected to a section failure
initiated by local buckling, while 60x45x15x1.6 and 90x45x15x1.6 mm studs failed by major
axis flexural buckling (Figs.11(a) and (b)). Fig. 10(a) shows that load ratio versus critical
outer hot flange temperature curve of 150x45x15x1.6 mm studs followed the trend of yield
strength reduction factor while those of 60x45x15x1.6 and 90x45x15x1.6 mm studs followed
the trend of elastic modulus reduction factor since major axis flexural buckling is influenced
by elastic modulus, while section yielding is influenced by yield strength. The critical outer
hot flange temperature curve of 150x45x15x1.6 mm stud did not coincide with the yield
strength reduction factor curve as cold flange temperatures and elastic modulus reduction
factors also affect its structural capacity.
There are significant differences in FRR when the stud depth was varied for LSF walls with
Configuration B (Fig. 9(b)), i.e. FRR values were 38, 40 and 60 minutes at a load ratio of 0.6
for the three HFC studs considered. However, FRR was about 80 minutes at a load ratio of
0.3, and below which FRRs were almost the same, irrespective of stud depth. Figs. 9(c) and
10(c) show that the load ratio versus FRR and critical hot flange temperature curves of the
cavity insulated walls (Configuration C) also significantly vary among them when the stud
depth was varied.
Overall, FRRs of LSF walls made of 150x45x15x1.6 mm studs were higher than for those
made of 90x45x15x1.6 and 60x45x15x1.6 mm studs with load ratios above 0.3, irrespective
of wall configuration type. The reasons are the lower thermal bowing deflections and the
section failure in the larger studs. Section failure mainly depends on yield strength whereas
major axis buckling failure mainly depends on elastic modulus. As seen in Fig. 10(a) the
yield strength reduction factor curve is above the elastic modulus curve for temperatures less
than 460oC, resulting in higher FRR of LSF walls made of 150x45x15x1.6 mm HFC sections
at higher load ratios. Beyond 460oC, the yield strength curve is below the elastic modulus
curve, leading to lower FRRs for 150x45x15x1.6 mm studs at lower load ratios.
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4. Effect of Steel Grade on the Fire Performance of LSF Walls
Elevated temperature mechanical property reduction factors play a vital role on the fire
performance of LSF walls. In this study, three steel types (Types A to C) were considered.
Their effects on the fire performance of LSF walls made of three LSF wall configurations
(Fig. 5) and 150x45x15x1.6 and 60x45x15x1.6 mm HFC section studs were investigated.
4.1. LSF Walls Made of 150x45x15x1.6 mm HFC Studs
Fig. 12(a) presents the load ratio versus FRR curves for Configuration A, which reveals the
large differences between the FRRs due to the use of studs made of different steel types. Fig.
13(a) presents the load ratio versus critical outer hot flange temperature curves, which follow
the trends of their respective yield strength reduction factors as 150x45x15x1.6 mm studs
failed by section yielding initiated by local buckling at both ambient and elevated
temperatures. As evident from Figs. 12(b) and (c) and 13(b) and (c), the load ratio versus
FRR and critical hot flange temperature curves of LSF walls with Configurations B and C are
also considerably influenced by the different steel types used.
4.2. LSF Walls Made of 60x45x15x1.6 mm HFC Studs
The FRR of LSF walls with Configuration A changes significantly when studs made of
different steels were used (Fig. 14(a)). As seen in Fig. 15(a), the load ratio versus critical
outer hot flange temperature curves were almost parallel to their respective elastic modulus
reduction factor curves as 60x45x15x15x1.6 mm studs failed by major axis flexural buckling.
Despite the fact that Types A and B steels have the same elevated temperature elastic
modulus reduction factors, their load ratio versus critical outer hot flange temperature curves
were slightly different. This is because the failures were also influenced by yield strength.
The trends in the load ratio versus FRR and critical hot flange temperature curves of LSF
walls with Configurations B and C were similar to those with Configuration A (Figs. 14(b)
and (c) and 15(c) and (d)).
Overall, the large differences among the FRR of LSF walls due to the use of studs made of
different steel types illustrate the significant influence of elevated temperature mechanical
property reduction factors on the fire performance of LSF walls. This also demonstrates the
need to use accurate mechanical property reduction factors in predicting the FRR of LSF
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walls using FEA or design models. This study has shown that using steels with higher
elevated temperature mechanical property reduction factors such as Type C steel would
significantly enhance the fire performance of LSF walls. This is likely to be more cost
effective than using additional plasterboard layers to increase the FRR of LSF walls.
5. Effect of Stud Thickness on the Fire Performance of LSF Walls
Previously Feng and Wang [23] conducted fire tests of LSF walls made of LCS studs of
different thicknesses (100x56x15x2 and 100x54x15x1.2 mm), and found that LSF walls
made of thicker studs had higher FRRs. They stated this was due to the differences in
temperature effects on steel sections with different thicknesses such as the changes in
effective widths of steel plates of varying thickness being different at the same temperature.
Contradicting this finding, Gunalan and Mahendran [24] found that LSF walls made of
thinner studs had higher FRRs due to the higher bending moment developed by the higher
load and the thermal bowing deflection in thicker studs. Their justification is questionable as
although the induced bending actions are higher in thicker studs, their ultimate bending
capacities are also higher. Further, they used the same time-temperature profiles for LSF
walls made of studs with different thicknesses, which is not acceptable as they were different
according to Feng and Wang’s [23] and Kesawan and Mahendran’s [16] thermal studies.
Kesawan and Mahendran’s [16] SAFIR thermal analysis results for LSF walls with
Configuration C shown in Figs. 16(a) and (b) demonstrate the significant influence of steel
stud thickness on the temperature developments in outer hot and cold flanges (OHF, OCF).
This section investigates the performance of LSF walls made 1.0, 1.6 and 2.5 mm thick G500
steel studs under standard fire conditions [7]. The web and flange sizes were maintained as
150 and 45 mm, respectively. Fig. 17(a) shows that the fire performance of LSF walls made
of 2.5 mm thick stud is the best for wall Configuration A, followed by those made of 1.6 and
1.0 mm thick studs. This is because the hot flange temperature development of thicker studs
is slow (Fig. 16(a)) with reduced local buckling effects. The reduced fire performance of LSF
walls with thinner studs is also evident from the load ratio versus critical hot flange
temperature curves (Fig. 18(a)) where the curves of 1.6 and 2.5 mm thick studs almost
coincided while that of 1.0 mm thick stud is below them. As seen in Figs. 17(b) and 18(b),
the load ratio versus FRR and critical hot flange temperature curves of LSF walls with
Configuration B have similar trends observed with Configuration A.
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Unlike in LSF walls with Configurations A and B, the influence of steel thickness on the
structural performance of LSF walls under fire conditions is more significant in cavity
insulated LSF walls – Configuration C (Fig. 17(c)). This was because the effect of stud
thickness on the temperature development of steel studs is higher in cavity insulated LSF
walls in comparison to uninsulated LSF walls [16]. Further, the reduction in cold flange
temperatures with increasing stud thickness (Fig. 16(b)) did not exert considerable influence
on the FRR of LSF walls. Fig. 18(c) presents the load ratio versus critical hot flange
temperature curve for LSF walls with Configuration C.
Overall, the steel stud thickness significantly affects the fire performance of LSF walls with
or without cavity insulation. LSF walls made of thicker studs have higher FRRs due to the
slower hot flange temperature developments and reduced local buckling effects in their plate
elements.
6. Effect of Realistic Design Fire Curves on the Fire Performance of LSF Walls
The ability of the standard fire curve to represent the modern day fires has become
questionable because of the increased severity of damage to buildings by accidental fires [25-
27]. Eurocode 3 Part 1-2 [19] - Annex A gives simple mathematical formulae to derive the
time-temperature relationships under fire conditions based on the available fuel load,
ventilation conditions and thermal inertia of walls. Ariyanayagam and Mahendran [28] found
that realistic design fire curves are more critical for single plasterboard lined LSF walls and
used eight different Eurocode parametric curves for them. Two of their curves (Fig. 19(a));
one prolonged (EU1) and one rapid growth (EU6) fire curves, were chosen in this study. The
parameters used to derive these fire curves can be found in Ariyanayagam and Mahendran
[28]. Thermal analyses in Kesawan and Mahendran [16] gave the outer hot and cold flange
time-temperature histories (Fig. 19(b)) of LSF walls made of 150x45x15x1.6 mm G500 HFC
section studs exposed to EU1 and EU6 curves, which were used in this study.
As seen in Fig. 19(c), FRR of LSF walls significantly reduced when EU1 and EU6 curves
were used, i.e., for a load ratio of 0.4, the FRR decreased from 73 minutes in the standard fire
to 35 and 58 minutes when the studs were exposed to EU1 and EU6 time-temperature curves,
respectively. This is not acceptable if real fires in buildings follow the EU1 and EU6 time-
temperature curves instead of the standard fire curve. As seen in Fig. 19(d) the critical outer
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hot flange temperature versus load ratio curves coincided because the difference between the
hot and cold flange temperatures was nearly the same for LSF walls exposed to different fire
curves when the wall configuration and the stud sizes were the same (Fig. 19(b)).
7. Effect of Wall Configurations on the Fire Performance of LSF Walls
This section evaluates the performance of LSF walls with different wall configurations and
subjected standard fire conditions [7]. Five different wall configurations made of
150x45x15x1.6 mm HFC section studs (G500 steel) were considered (Fig.20). It was found
that LSF walls lined with three plasterboard layers has the highest FRR than the other wall
configurations (Fig. 21), and they should be used when higher fire performance is required.
The fire performance of cavity insulated LSF walls was poor in comparison to that of
uninsulated LSF walls. This was because the cavity insulation acted as a heat barrier resulting
in rapid hot flange temperatures rise, and higher induced bending actions due to the larger
differences between the hot and cold flange temperatures. The externally insulated LSF walls
(Fig. 20(d)) proposed by Kolarkar and Mahendran [29] performed better than the uninsulated
and cavity insulated LSF walls (185 minutes at a load ratio of 0.4 in comparison with 165 and
125 mins for others - Fig. 21). These results show that the fire performance of LSF walls was
significantly affected by the types of wall configuration including the number of plasterboard
layers, and the provision of insulation and its location.
8. Effect of Connectivity between Plasterboards and Steel Studs
LSF wall studs are connected to plasterboards using buggle head screws. These connections
provide lateral supports to the LSF wall studs, by which their torsional and minor axis
buckling deformations are eliminated. This significantly increases the structural capacity of
LSF wall studs. The screws can penetrate either through both the outer and inner flanges or
only through the outer flanges of HFC studs (Fig. 1). The effects of these connections were
investigated in this study. LSF walls with Configuration A and made of 150x45x15x1.6 mm
HFC section studs (G500 steel) were considered. Fig. 22 shows that the load ratio versus
FRR curves were about the same for a given wall configuration, and thus indicate that the
screw connectivity does not influence the fire performance of LSF wall studs. However, the
improved connectivity can reduce the level of plasterboard fall-off in LSF walls and fire tests
are needed to verify this.
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9. Fire Performance of LSF Walls Made of Studs with Different Section Geometry
Generally LCS studs are used in LSF walls, but recently Kesawan and Mahendran [1]
proposed to use the HFC section studs. This section investigates whether the structural
efficiency of the HFC section studs is higher than the LCS studs under fire conditions.
150x45x15x1.6 mm and 90x45x15x1.6 mm G500 HFC section and LCS studs (Type A steel)
were considered (Fig. 23). Here, the section depth, flange width and thickness of LCSs and
HFC sections were the same. Temperature development in these studs was assumed to be the
same as Kesawan and Mahendran [16] found that stud section profiles did not affect the
thermal performances under fire conditions, if they are of the same thickness and flange
width. Figs. 24(a) to (d) demonstrate that there is negligible difference between the FRR of
LSF walls made of HFC section and LCS studs. Therefore using HFC studs does not provide
an advantage over other sections with similar overall sizes and thickness in terms of retaining
the ambient temperature load carrying capacity of LSF wall studs exposed to fire conditions.
10. Critical Temperature Method to Predict the FRR of LSF Walls
The critical temperature method to determine the FRR of structural elements in buildings is
popular due to its simplicity. Eurocode 3 Part 1.2 [19] proposes a critical temperature of
350oC for cold-formed steel members while Lawson [30] and Kolarkar [31] recommended
improved critical temperature methods for LCS sections subject to non-uniform temperature
distributions when used in LSF walls exposed to fire on one side. Using detailed parametric
studies, Gunalan and Mahendran [32] proposed new critical hot flange temperature equations
for seven different wall configurations made of 90x40x15x1.15 mm LCS studs and exposed
to the standard fire curve [7] while Ariyanayagam and Mahendran [33] proposed improved
critical temperature equations for LSF walls exposed to realistic design fire curves. However,
they are limited to walls made of 90x40x15x1.15 mm LCS studs, and are not the same for
different wall configurations. Therefore an investigation was undertaken to evaluate the
influence of temperature distributions on the fire performance of LSF walls made of three
HFC studs, 150x40x15x1.15, 90x40x15x1.1.5 and 60x40x15x1.15 mm of Type A (G500)
steel. Gunalan and Mahendran [11] also used studs of 1.15 mm thickness and 40 mm flange
width. This study aims to answer two questions,
1. Whether the critical hot flange temperature changes with temperature distributions
2. Whether the critical hot flange temperature changes with stud sizes for similar
temperature distributions
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Both uniform and non-uniform temperature distributions (0, 100, 200, 300 and 400oC
differences between the outer hot and cold flanges) were considered as shown in Fig. 25, and
the results are shown in Figs. 26(a) to (e) as load ratio versus critical hot flange temperature
curves. These figures show that the critical hot flange temperature for a particular load ratio
varied with the stud sections used. Even the critical hot flange temperature method is not an
appropriate solution for studs subjected to uniform temperature distributions (Fig. 26(a)). As
seen in Figs. 27(a) to (c), the load ratio versus critical hot flange temperature curves were
dissimilar for the same stud section exposed to different temperature distributions. Therefore
the applicability of the critical hot flange temperature method proposed by previous
researchers is questionable, and should be restricted to the wall configurations and the stud
sizes used by them. In conclusion, it is not possible to propose critical hot flange temperature
equations for HFC section studs subject to non-uniform temperature distributions under fire
conditions.
The applicability of the critical average temperature method (average of hot and cold flange
temperatures) to predict the FRR of LSF walls was also evaluated. Figs. 28(a) to (c) present
the load ratio versus critical average temperature curves for the three HFC studs. The trends
in these curves are the same as in the load ratio versus critical hot flange temperature curves.
The critical average temperature for a particular load ratio varied with the temperature
distribution patterns. The higher the temperature difference between hot and cold flanges, the
lower the load ratio for a particular average temperature. The load ratio versus critical
average temperature curves of different sized studs for a temperature difference of 200oC in
Fig. 28(d) are dissimilar for such sections. This demonstrates that the average critical
temperature method is also not suitable to determine the FRR of LSF walls made of HFC
sections. However, the load ratio versus critical temperature curves produced for many HFC
sections subject to varying temperature distributions given in Figs. 26 to 28 can be used to
predict the FRR of LSF walls if the time-temperature history across the LSF wall is known.
7. Conclusions
This paper has presented an extensive finite element analysis based parametric study on the
structural fire performance of LSF walls under fire conditions. The effects of important
parameters such as stud sizes, stud profiles, wall configurations and elevated temperature
mechanical property reduction factors on the FRR of LSF walls were fully investigated.
14
Using the results obtained from this study, the load ratio versus FRR and critical hot flange
temperature curves were produced for all the different LSF walls considered in this study.
The results showed that the use of HFC studs with larger depth and thickness improved the
FRR of LSF walls while demonstrating the importance of using studs made of steels with
improved elevated temperature mechanical properties. Uninsulated and externally insulated
LSF walls provided enhanced fire performance than cavity insulated LSF walls. The
enhanced level of plasterboard to stud connectivity in HFC studs did not seem to affect the
structural fire performance of LSF walls, while the use of HFC studs did not provide an
advantage over other stud sections with similar overall sizes and thickness in terms of
retaining the ambient temperature load carrying capacity of LSF wall studs exposed to fire
conditions. FRR of LSF walls was significantly affected when realistic design fire curves
were used instead of the standard fire curve, demonstrating the importance of using
appropriate realistic design fire curves in design.
Both the critical hot and average flange temperature methods were found to be unsuitable to
predict the FRR of LSF walls made of HFC studs. However, the load ratio versus FRR and
critical hot flange temperature curves of many LSF wall systems produced in this study can
be used in fire design. The developed comprehensive fire performance data would facilitate
the development of LSF walls with enhanced fire performance and importantly it would
facilitate and advance the successful applications of HFC section studs in LSF walls.
Moreover, it would pave the path for developing a performed based approach for fire design
of LSF walls.
Acknowledgements
The authors would like to thank Queensland University of Technology (QUT) and Australian
Research Council for providing the financial support to conduct this research project and
QUT for providing the high performance computing facilities.
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1
(a) LCS (b) Welded HFC section (c) Screw/rivet fastened HFC section
Figure 1: LSF Wall Studs
Figure 2: Full Scale Fire Test of LSF Walls
Plasterboards
Screw Screw
Test Wall Panel
Loading
System
Supporting
Frame
Screw/rivet
fastening
2
Figure 3: Boundary and Loading Conditions
Figure 4: Temperature Distribution across LSF Wall Studs
Restrained
DOF ‘234’
Restrained
DOF ‘1234’
Restrained
DOF ‘3’
Restrained
DOF ‘3’
Restrained DOF ‘3’ at 300 mm
(Lateral restraints provided by the plasterboards)
300 mm
Y
Z
X
Load
Ambient Side
Fire Side
Linear temperature
distribution
Uniform
temperature
distribution
3
Figure 5: LSF Wall Configurations (Kesawan and Mahendran, 2015c)
Figure 6: Initial Geometric Imperfections
(a) Non-insulated LSF walls with dual plasterboard layers (Configuration A)
(b) Non-insulated LSF walls with single plasterboard layer (Configuration B)
Stud A Stud B Stud C Stud D
Fire Side
(FS)
Ambient Side
(AS)
Outer Hot Flange (OHF)
Inner Hot Flange (IHF)
Mid-web
Inner Cold Flange (ICF) Outer Cold Flange (OCF)
Stud A Stud B Stud C Stud D
l
(b) Local Imperfection
b/150 b b/150
(a) Global Imperfection
(c) Insulated LSF walls with dual plasterboard layers (Configuration C)
50 mm thick insulation Stud A Stud B Stud C Stud D
4
Figure 7: HFC Section Studs with Different Web Depths
Figure 8: Effect of Stud Depth on the Time-Temperature Profiles of LSF Walls
(Configuration A)
0
100
200
300
400
500
600
700
0 20 40 60 80 100 120 140 160 180 200 220 240 260
Tem
pe
ratu
re (
oC
)
Time (minutes)
OHF-60 mm OCF-60 mm OHF-90 mmOCF-90 mm OHF-150 mm OCF-150 mm
5
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 9: Load Ratio versus FRR Curves for LSF Walls with HFC Section
Studs of Varying Depths (60, 90 and 150 mm)
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200 220 240
Load
Rat
io
Failure Time (minutes) 60 mm 90 mm 150 mm
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140
Load
Rat
io
60 mm 90 mm 150 mm
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200
Load
Rat
io
Time (minutes) 60 mm 90 mm 150 mm
6
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 10: Load Ratio versus Critical Outer Hot Flange Temperature Curves
for LSF Walls with HFC Section Studs of Varying Depths (60, 90 and 150 mm)
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700
Load
Rat
io
Outer Hot Flange Temperature (oC) 60 mm 90 mm 150 mm
fy reduction factor E reduction factor
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800
Load
Rat
io
Outer Hot Flange Temperature (oC) 60 mm 90 mm 150 mm
fy reduction factor E reduction factor
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) 60 mm 90 mm 150 mmfy reduction factor E reduction factor
7
(a) 90x45x15x1.6 mm
(a) 150x45x15x1.6 mm
Figure 11: Failure Modes of HFC Section Studs in LSF Walls (Configuration A)
Local buckling waves
Section Failure
Major axis
flexural buckling
8
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 12: Load Ratio versus FRR Curves for LSF Walls with 150x45x15x1.6
mm HFC Section Studs Made of Different Steel Types
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200 220 240
Load
Rat
io
Failure Time (minutes)
Type A Type B Type C
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140
Load
Rat
io
Failure Time (minutes)
Type A Type B Type C
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200
Load
Rat
io
Failure Time (minutes)
Type A Type B Type C
9
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 13: Load Ratio versus Critical Outer Hot Flange Temperature Curves
for LSF Walls with 150x45x15x1.6 mm HFC Section Studs Made of Different
Steel Types
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600
Load
Rat
io
Outer Hot Flange Temperature (oC) Type A Type BType C Type A - fy reduction factorType B - fy reduction factor Type C -fy reduction factor
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 200 400 600 800 1000
Load
Rat
io
Outer Hot Flange Temperature (oC) Type A Type BType C Type A - fy reduction factorType B - fy reduction factor Type C - fy reduction factor
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 200 400 600 800 1000
Load
Rat
io
Outer Hot Flange Temperature (oC) Type A Type B
Type C Type A - fy reduction factor
Type B - fy reduction factor Type C - fy reduction factor
10
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 14: Load Ratio versus FRR Curves for LSF Walls with 60x45x15x1.6
mm Studs Made of Different Steel Types
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200 220 240
Load
Rat
io
Failure Time (minutes) Type A Type B Type C
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140
Load
Rat
io
Failure Time (minutes) Type A Type B Type C
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200
Load
Rat
io
Failure Time (minutes) Type A Type B Type C
118
11
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 15: Load Ratio versus Critical Outer Hot Flange Temperature Curves
for LSF Walls with 60x45x15x1.6 mm HFC Section Studs Made of Different
Steel Types
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700
Load
Rat
io
Outer Hot Flange Temperature (oC) Type A Type BType C Types A & B- E reduction factorType C - E reduction factor
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700
Load
Rat
io
Outer Hot Flange Temperature (oC) Type A Type BType C Types A and B - E reduction factorType C - E reduction factor
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700
Load
Rat
io
Outer Hot Flange Temperature (oC) Type A Type BType C Types A & B - E reduction factorType C - E reduction factor
12
(a) Outer hot flange
(b) Outer cold flange
Figure 16: Time-Temperature Profiles of LSF Walls with Configuration C
0
100
200
300
400
500
600
700
800
900
0 20 40 60 80 100 120 140 160 180 200 220 240
Tem
pe
ratu
re (
oC
)
Time (minutes) OHF-1 mm OHF-1.6 mm OHF-2.5 mm
0
50
100
150
200
250
300
350
400
450
0 20 40 60 80 100 120 140 160 180 200 220 240 260
Tem
pe
ratu
re (
oC
)
Time (minutes)
OCF-1 mm OCF-1.6 mm OCF-2.5 mm
13
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 17: Load Ratio versus FRR Curves for LSF Walls with Studs of Varying
Thicknesses (1, 1.6 and 2.5 mm)
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200 220 240
Load
Rat
io
Failure Time (minutes) 1mm 1.6 mm 2.5 mm
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140
Load
Rat
io
Failure Time (minutes) 1 mm 1.6 mm 2.5 mm
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200
Load
Rat
io
Failure Time (minutes) 1 mm 1.6 mm 2.5 mm
14
(a) Wall Configuration A
(b) Wall Configuration B
(c) Wall Configuration C
Figure 18: Load Ratio versus Critical Outer Hot Flange Temperature Curves
for LSF Walls with Studs of Varying Thicknesses (1, 1.6 and 2.5 mm)
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800
Load
Rat
io
Outer Hot Flange Temperature (oC) 1 mm 1.6 mm 2.5 mm
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800
Load
Rat
io
Outer Hot Flange Temperature (oC) 1 mm 1.6 mm 2.5 mm
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800
Load
Rat
io
Outer Hot Flange Temperature (oC) 1 mm 1.6 mm 2.5 mm
15
(a) Eurocode Parametric and Standard Fire Curves
(b) Outer Hot and Cold Flange Time-Temperature Profiles
(c) Load Ratio versus FRR Curves
0
200
400
600
800
1000
1200
1400
0 30 60 90 120 150 180 210 240 270
Tem
pe
ratu
re (
oC
)
Time (minutes) EU1 EU6 Standard Fire Curve
0
100
200
300
400
500
600
700
800
0 20 40 60 80 100 120 140 160 180 200 220 240
Tem
pe
ratu
re (
oC
)
Time (minutes) EU1-OHF EU1-OCFEU6-OHF EU6-OCFStandard Curve-OHF Standard Curve-OCF
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 10 20 30 40 50 60 70 80 90 100 110 120 130 140
Load
Rat
io
Failure Time (minutes) Standard Curve EU1 EU6
73 58 35
16
(d) Load Ratio versus Critical Hot Flange Temperature Curves
Figure 19: Uninsulated and Single Plasterboard Lined LSF Walls Exposed to
Standard and Realistic Design Fire Curves
Figure 20: Different Wall Configurations
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800
Load
Rat
io
Outer Hot Flange Temperature (oC)) Standard Curve EU1 EU6
(a) Uninsulated LSF Walls Lined with Single Plasterboard
(b) Uninsulated LSF Walls Lined with Dual Plasterboards
(c) Uninsulated LSF Walls Lined with Three Plasterboards
(e) Cavity Insulated LSF Walls Lined with Dual Plasterboards
(d) Externally Insulated LSF Walls Lined with Dual Plasterboards Figure 20:
17
Figure 21: Load Ratio versus FRR Curves for LSF Walls with Different Wall
Configurations
Figure 22: Load Ratio versus FRR Curves for LSF Walls Made of
150x45x15x1.6 mm Studs with Different Stud to Plasterboard Connections
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 50 100 150 200 250 300 350
Load
Rat
io
Failure Time (minutes)
Uninsulated - 2Pb Uninsulated-Pb Insulated-Cavity-2Pb
Uninsulated-3Pb Insulated-External-2Pb
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200 220 240
Load
Rat
io
Failure Time (minutes)
2Pb - Inner and Outer Pb - Inner and Outer
2Pb - Ins - Inner and Outer 2Pb- Outer
Pb - Outer 2Pb - Ins - Outer
125
165
185
18
Figure 23: Different Section Profiles
(a) Configuration A with 150x45x15x1.6 mm studs
(b) Configuration B with 150x45x15x1.6 mm studs
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200 220 240
Load
Rat
io
Failure Time (minutes)
Lipped Channel Section Hollow Flange Section
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140
Load
Rat
io
Failure Time (minutes)
Lipped Channel Section Hollow Flange Section
(a) (b) (c) (d)
19
(c) Configuration C with 150x45x15x1.6 mm studs
(d) Configuration A with 90x45x15x1.6 mm studs
Figure 24: Load Ratio versus FRR of LSF Walls with G500 Steel Studs
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200
Load
Rat
io
Failure Time (minutes)
Lipped Channel Section Hollow Flange Section
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 20 40 60 80 100 120 140 160 180 200
Load
Rat
io
Failure Time (minutes)
Lipped Channel Section Hollow Flange Section
20
Figure 25: Uniform and Non-uniform Temperature Distribution Patterns
21
(a) Uniform Temperature Distribution
(b) 1000C Difference
(c) 2000C Difference
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) 60x40x15x1.15 90x40x15x1.15 150x40x15x1.15
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) 60x40x15x1.15 90x40x15x1.15 150x40x15x1.15
0.00
0.10
0.20
0.30
0.40
0.50
0.60
0.70
0.80
0.90
1.00
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) 60x40x15x1.15 90x40x15x1.15 150x40x15x1.15
22
(d) 3000C Difference
(e) 4000C Difference
Figure 26: Load Ratio versus Critical Outer Hot Flange Temperature Curves
for Studs Subject to Non-Uniform Temperature Distributions
0.00
0.10
0.20
0.30
0.40
0.50
0.60
0.70
0.80
0.90
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) 60x40x15x1.15 90x40x15x1.15 150x40x15x1.15
0.00
0.10
0.20
0.30
0.40
0.50
0.60
0.70
0.80
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) 60x40x15x1.6 90x40x15x1.6 150x40x15x1.6
23
(a) 60x40x15x1.15 mm Studs
(b) 90x40x15x1.15 mm Studs
(c) 150x40x15x1.15 mm Studs
Figure 27: Load Ratio versus Critical Outer Hot Flange Temperature Curves
for Different HFC Stud Sections
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) Uniform 100 200 300 400
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) Uniform 100 200 300 400
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Outer Hot Flange Temperature (oC) Uniform 100 200 300 400
24
(a) 60x40x15x1.15 mm Studs
(b) 90x40x15x1.15 mm Studs
(c) 150x40x15x1.15 mm Studs
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Average Temperature (oC) Uniform 100 200 300 400
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Average Temperature (oC) 0 100 200 300 400
0.00
0.20
0.40
0.60
0.80
1.00
1.20
0 100 200 300 400 500 600 700 800 900
Load
Rat
io
Average Temperature (oC)
Uniform 100 200 300 400
25
(d) 2000C Difference
Figure 28: Load Ratio versus Critical Average Temperature Curves
0.000.100.200.300.400.500.600.700.800.901.00
0 100 200 300 400 500 600 700 800
Load
Rat
io
Average Temperature (oC)
60x40x15x1.15 90x40x15x1.15 150x40x15x1.15
1
Table 1: Fire Test and FEA Results (Kesawan and Mahendran, 2015)
Note: All the tests were conducted under standard fire conditions
Wall Configurations Load
Ratio
Fire Tests-
FRR (mins.)
FEA –
FRR (mins.)
Transient
State
Steady
State
Test 1
0.4 180 183 184
Test 2
0.2 205 208 209
Test 3
0.2 136 125 127
Test 4
0.2 182 180 180
Test 5
0.6 138 136 137
2
Table 2: Elevated Temperature Yield Strength Reduction Factors
Temperature
(oC)
Dolamune Kankanamge and
Mahendran (2011) Eurocode 3 Part
1.2 (ECS, 2005) High Strength Low Strength
20 1.000 1.000 1.000
50 0.995 0.985 1.000
100 0.986 0.960 1.000
150 0.977 0.935 0.945
200 0.968 0.910 0.890
250 0.959 0.771 0.835
300 0.950 0.658 0.780
350 0.810 0.561 0.715
400 0.670 0.478 0.650
450 0.530 0.404 0.590
500 0.390 0.337 0.530
550 0.250 0.277 0.415
600 0.110 0.222 0.300
650 0.090 0.171 0.215
700 0.070 0.124 0.130
750 0.050 0.080 0.100
800 0.030 0.039 0.070
Table 3: Elevated Temperature Elastic Modulus Reduction Factors
Temperature
(oC)
Dolamune Kankanamge
and Mahendran (2011)
Eurocode 3 Part
1.2 (ECS, 2005)
20 1.000 1.000
50 0.975 1.000
100 0.933 1.000
150 0.892 0.950
200 0.850 0.900
250 0.783 0.850
300 0.715 0.800
350 0.648 0.750
400 0.580 0.700
450 0.513 0.650
500 0.445 0.600
550 0.378 0.455
600 0.310 0.310
650 0.243 0.220
700 0.175 0.130
750 0.108 0.110
800 0.040 0.090