Developing a performance factor for fire rated boards used in LSF wall systems

Fire Safety Journal 109 (2019) 102872

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Developing a performance factor for fire rated boards used in LSF wall systems

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Maneesha Tharindi Dodangoda, Mahen Mahendran∗, Poologanathan Keerthan, Ray L. Frost Queensland University of Technology (QUT), Brisbane, QLD, 4000, Australia

ARTICLE INFO

ABSTRACT

Keywords: Gypsum plasterboards Thermo-physical properties Light-gauge steel frame wall systems Finite element heat transfer analysis Fire performance factor

Recently unclassified boards with varying composition and more notably with ambiguous thermal properties are being used in buildings. The fire performance of Light gauge Steel Frame (LSF) wall systems mainly depends on the protective plasterboard linings used and thus fire safety concerns are being raised. Therefore this study was aimed at setting minimum standards for fire-resistant grade plasterboards used in LSF wall applications based on material characterization, thermo-physical properties and finite element heat transfer modelling. The most commonly used fire protective board, gypsum plasterboard was used to address this critical issue. Thermophysical properties of three gypsum plasterboards manufactured in Australia were measured and compared in relation to their chemical composition in the first phase of the study. Fire resistance levels (FRLs) of LSF wall systems lined with these three plasterboards were determined using 3-D FE heat transfer models of LSF wall systems developed and validated in this study. The paper proposes a “k-factor” based on thermo-physical properties and thickness of plasterboard, capable of giving an overall measure of the fire performance of plasterboard lined LSF walls. Standard k-factor profiles were established for non-load bearing LSF wall systems with FRLs of 60, 120, 180 and 240 min by considering the correlation between the time-temperature profiles from numerical analyses and calculated k-factors. These proposed standard k-factor profiles give an overall measure of the fire performance of plasterboards. This paper presents the details of this study and the results.

1. Introduction Buildings must be designed and constructed to meet acceptable standards of structural adequacy, safety, health and services. Fire safety is one of the most important considerations of building construction industry. In the National Construction Code (NCC) of Australia [1] fireresistant grade plasterboards are identified as suitable fire protective covering for building elements. These plasterboards are popularly used with light gauge steel framed (LSF) wall systems, and are considered as fire rated barriers. The NCC specifies certain Fire Resistance Levels (FRL) for construction wall elements for them to be able to provide fire compartmentation. The required FRLs can be 30, 60, 90, 120, 180 or 240 min for both load bearing and non-load bearing walls depending on the class of building. LSF wall systems are used with or without cavity insulation depending on the energy requirement and performance of the building. The steel frame of LSF wall systems is made of thin-walled coldformed lipped channel section (LCS) studs and unlipped channel section tracks. When exposed to fire conditions, these thin-walled steel stud sections heat up rapidly and reach their failure temperatures quickly. It

∗

will eventually lead to structural instability of load bearing LSF wall systems and thus of the building. Therefore, fire resistance of load bearing LSF wall systems mainly depend on the protective linings in use, i.e. fire-resistant grade boards, which keep the steel stud temperatures from reaching their limiting temperatures. Fire rated gypsum plasterboards are the most commonly used type of boards as protective lining for LSF wall systems. The FRL of load bearing LSF wall systems protected with single and double layers of fire rated 16 mm gypsum plasterboard is approximately 60 and 120 min, respectively [2–4]. Each additional layer of 16 mm gypsum plasterboard increases the FRL of load bearing LSF wall systems by about 60 min [2–4], while it is about 90 min for non-load bearing LSF walls. In recent times, unclassified boards with varying composition and more notably with ambiguous thermal properties are increasingly appearing in the building construction market. These boards claim improved fire performance and very high FRLs in conjunction with LSF wall applications. Full-scale standard fire tests of LSF wall systems lined with these boards have to be conducted based on AS 1530.4 [5] to satisfy the FRL requirements in Ref. [1]. However, expert opinions or over-simplified tests are often used to estimate their FRLs instead of full

Corresponding author. E-mail address: [email protected] (M. Mahendran).

https://doi.org/10.1016/j.firesaf.2019.102872 Received 8 February 2017; Received in revised form 13 October 2018; Accepted 8 September 2019 Available online 11 September 2019 0379-7112/ © 2019 Elsevier Ltd. All rights reserved.

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scale standard fire tests to AS 1530.4. Since full scale fire tests are expensive and time consuming, some fire protective board suppliers claim their boards to be equivalent to other gypsum plasterboards for whom standard fire test results are available. Hence there is a need to implement minimum standards for fire protective boards in order to ensure appropriate boards are used as the fire-protective covering for building elements. Therefore this study proposes a minimum standard for fire-resistant grade boards based on small scale thermo-physical property tests. There is also a need to study the effect of these thermophysical properties on the FRL of LSF walls and to determine whether an approximate method can be developed to estimate the FRL based on thermal properties. Hence a study was undertaken based on experimental material characterization and finite element thermal modelling of the most commonly used fire protective board, fire resistant grade gypsum plasterboards, to address these critical issues on fire safety design. This study provides a better understanding of gypsum plasterboards’ chemical and physical properties at elevated temperatures and fire performance of board panels exposed to ISO 834 [6] standard fire (identical fire curve in AS 1530.4). Fire-resistant grade gypsum plasterboards from three different manufacturers in Australia were considered in this study.

Supplementary to gypsum, the composition of commercial plasterboards consists of three other categories of constituents. These materials are accountable for adjusting shrinkage, maintaining integrity, and providing low thermal conductivity of the end product exposed to fire. Vermiculite, glass fibres and fillers are used respectively under the above categories. When heated to above 300 °C vermiculite, a natural occurring mineral of hydro-mica group [12] undergoes a significant expansion. This exfoliation compensates for the volume loss from gypsum dehydration and acts as a shrinkage adjusting element [13]. Fire-resistant grade gypsum plasterboards are typically reinforced with glass fibres. Glass fibres act as a bridging material during heat induced chemical reactions of the gypsum core where it loses part of its mass. Thus glass fibres help in maintaining the integrity of gypsum core up to its melting temperatures above 800 °C. After that point plasterboards undergo extensive ablation and cracking. Thomas' [9] anecdotal evidence of fire test observations verified this. In addition, glass fibres assist in increasing the mechanical properties of plasterboards. Different fillers are also present in small quantities depending on manufacturers’ specifications. Typically these fillers are low density and low thermal conductive materials to make the boards lighter and also to make them less thermal conductive.

2. Gypsum plasterboards

3. Experimental studies

Fire-resistant grade gypsum plasterboards are manufactured by sandwiching a dense gypsum core between heavy duty recycled papers. The main constituent of this gypsum core, gypsum, is a naturally occurring soft sulphate mineral, and is commonly found in calcareous sedimentary rock deposits. It is chemically acknowledged as Calcium Sulphate Dihydrate (CaSO4·2H2O). The percentage of pure gypsum (i.e. CaSO4·2H2O) within can vary for different plasterboards depending on the manufacturer. Wakili and Hugi's [7] study of fire behaviour of four different European gypsum plasterboards showed it to be between 60 and 100%. This pure gypsum contains approximately 20.9% chemically bound water. Additionally, about 3–4% free moisture content [8,9] is present within the pores of gypsum core depending on the ambient temperature and relative humidity. When gypsum plasterboards are exposed to fire, energy in the form of heat is absorbed to evaporate this free and crystalline water contents. This heat absorption delays the heat transfer across the plasterboard thickness. The chemical reaction involving the release of water molecule from the reacting molecule is known as dehydration, and it occurs in two phases. At first, it loses ¾ of its chemically bound water and results in the powdery material called, Plaster of Paris (CaSO4·½H2O – Calcium Sulphate Hemi-hydrate) [10] (Reaction 1). In the second dehydration process remaining ¼ of chemically bound water is released from Calcium Sulphate Hemi-hydrate and result in CaSO4 (III) – Calcium Sulphate Anhydrite III [10] (Reaction 2). The starting and ending temperatures for these two reactions can span in the temperature regions of 80–140 °C and 110–220 °C, respectively, depending on the heating rate [10].

3.1. Materials

Reaction 2:CaSO4 1 2 H2 O

Reaction 4:CaSO4 (II)

3.2. Methods – chemical composition characterization

CaSO4 1 2 H2 O + 3 2 H2 O

Reaction 1:CaSO4 2H2 O

Reaction 3:CaSO4 (III)

In the experimental studies fire-resistant grade gypsum plasterboards from three Australian manufacturers (Boards 1, 2 and 3) were studied for chemical composition, thermo-physical properties and fire performance. In Australia, fire-resistant grade gypsum plasterboards are available in two thicknesses, 13 mm and 16 mm with pink colour paper liner. The 16 mm thick plasterboards were investigated in this study as it is the standard plasterboard thickness for fire-resistant grade gypsum plasterboards with load bearing LSF wall systems to get FRL of 120 min. For all the chemical and thermo-physical characterization tests, the paper liners of boards were removed first and samples for different tests were prepared from the gypsum core of each board. Firstly the surfaces of the cores of three fire-resistant grade gypsum plasterboards were visually inspected using light microscopy (LM). The images obtained of grinded surfaces and freshly cracked surfaces of gypsum core of three plasterboards are shown in Fig. 1(a) and (b), respectively. The physical appearance of Board 3 was different from other two boards with its comparatively larger size pores of 0.25–1.5 mm. The other two boards have smaller size pores of ≤0.25 mm. The glass fibres used for the purpose of reinforcing plasterboards were clearly visible on the freshly cracked surfaces of all three boards (Fig. 1(b)). The shrinkage adjusting material, vermiculite, was also clearly observable (black colour dots/ patches) in both Fig. 1(a) and (b).

3.2.1. X-ray diffraction (XRD) The powder X-ray diffraction (PXRD) analysis was undertaken in this study for gypsum plasterboard chemical composition characterization purposes. The gypsum plasterboard samples were used in powdered state by freshly grinding them. The specimens were prepared by weighing sub samples and adding corundum (Al2O3) as an internal standard at c.a.10 wt %. The X-ray diffraction patterns were collected with a PANalytical X'Pert Pro diffractometer using cobalt Kα radiation (40 kV, 40 mA). Samples were spun during data collection. The collected data was analysed using software JADE and Highscore Plus for phase identification, and TOPAS for quantitative phase analysis using the Rietveld method.

CaSO4 (III) + 1 2 H2 O

CaSO4 (II) CaSO4 (I)

At temperatures above 300 °C the resultant Calcium Sulphate Anhydrite III transforms to CaSO4 (II) – Calcium Sulphate Anhydrite II (Reaction 3). In this transformation, soluble hexagonal crystal structure irreversibly altered into an orthorhombic crystal structure of lower insoluble energy state [10,11]. The final reaction takes place at temperatures above 1180 °C where Sulphate Anhydrite II transforms into a cubic crystal structure of CaSO4 (I) – Calcium Sulphate Anhydrite I [10] (Reaction 4).

3.2.2. Scanning electron microscopy (SEM) The crystalline structure of the constituents of gypsum plasterboards 2

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Fig. 1. Visual images of the core of three gypsum plasterboards (a) ground surface, (b) freshly cracked surface.

was observed with FEI Quanta 200 environmental scanning electron microscopy (ESEM). The specimens were freshly cracked and mounted on to an aluminium SEM stub using double sided carbon tape. Images were obtained in low vacuum mode using a back scatter electron detector.

Fig. 3 show the crystalline structure of the constituents of gypsum plasterboards. It clearly shows the needle like structures of gypsum, plate like vermiculite and glass fibres.

3.3. Results and discussion – chemical composition characterization

3.4.1. Simultaneous thermal analyser (STA) Simultaneous thermal analysis (STA) of differential scanning calorimetry (DSC) and Thermogravimetric Analysis (TGA) measurements were conducted using NETZSCH STA 449F3 Jupiter. The instrument was calibrated for temperature and sensitivity using a melting point method and the materials used for calibration were In, Sn, Zn, Al, Ag and Au. The experiments were conducted in a purged environment using Nitrogen at a flow rate of 20 ml/min and with STC off. Three evacuations and refills of Nitrogen preceded each measurement to ensure an absolute purge environment before measurements. The temperature program for all the measurements was set to start from 50 °C and increase to 1200 °C at a steady rate of 20 °C/min. A 15 min isothermal segment was maintained at both 50 °C and 1200 °C. Platinum (Pt) crucibles lined with Alumina (Al2O3) liners and pin holed lids were used for all tests. Results were analysed using the NETZSCH Proteus® software. The DSC tests for specific heat variation were conducted according to the three step procedure specified in ASTM E1269 [14]. In this method, the difference in heat flow between the empty crucible and the reference material of known specific heat is recorded. The difference in heat flow between the test material and the empty crucible is also recorded. The specific heat is then calculated using a ratio between these two measurements and by incorporating the mass variation with temperature. The same empty reference crucible and sample crucible was used in all three steps of the tests and the reference crucible was not removed from the sample carrier. A 0.5 mm sapphire disc supplied by NETZSCH was used as the reference standard material of known specific heat. The gypsum plasterboard samples were freshly ground and each sample was measured for initial mass of ≈15 mg. The mass retention variations with temperature were obtained using TGA technique simultaneously.

3.4. Methods – thermo-physical characterization

Fire retarding property of gypsum plasterboards is mainly related to its main material component, Gypsum (Reactions 1 and 2). Therefore, the percentage of gypsum in plasterboards is important when comparing plasterboards for their thermo-physical properties and fire performance in LSF wall applications. However, the specific chemical identity and/or exact percentage of composition are not available for any of the commercially available gypsum plasterboards. Therefore, the powder X-ray diffraction (PXRD) analysis was undertaken in this study to find approximate percentages of Gypsum and other minerals in the three fire-resistant grade gypsum plasterboards considered. Table 1 shows the PXRD analysis results. Approximate percentages of Gypsum (CaSO4·2H2O) of Boards 1 and 3 are quite similar at values 83.9% and 84.4%, respectively, while it is much lower for board 2 at 65%. However, Board 2 has 8.8% of Bassanite (CaSO4·½H2O or 2CaSO4·H2O), which is a hemihydrate. When heated ≈84% of materials in Boards 1 and 3 will undergo both dehydration steps (Reactions 1 and 2). In the case of Board 2, approximately 65% of materials will undergo first dehydration step (Reaction 1) and approximately 73.8% of materials will undergo second dehydration step (Reaction 2). The Scanning electron microscopy (SEM) images (magnification 100 × ) of the three boards as shown in Fig. 2 confirmed the pore sizes and distribution pattern observed with LM. Boards 1 and 2 have large number of evenly spread smaller size pores (≤0.25 mm) while Board 3 has comparatively large scale pores (0.25–1.5 mm). SEM images in Table 1 Mineral composition from PXRD analysis. Components

Board 1

Board 2

Board 3

Quartz Aragonite Anhydrite Bassanite Gypsum Vermiculite Sepiolite Non-diffracting/unidentified

0.6 2.4 1.5 2.2 83.9 1.7 5.7 2.0

1.2 3.3 1.3 8.8 65.0 0.8 7.0 12.7

1.4 3.4 1.9 1.2 84.4

3.4.2. Dilatometer (DIL) Linear thermal expansion variations with temperature were measured using a NETZSCH DIL 402C dilatometer. The instrument measures the expansion or shrinkage of test samples during thermal treatment. The sample holder and push rod of the instrument were Alumina (Al2O3). The instrument was calibrated for temperature using a melting

7.8

3

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Fig. 2. SEM images (pore sizes and distribution) (a) Board 1 (b) Board 2 (c) Board 3.

point method and the standard materials were In, Sn, Zn, Al, Ag and Au. The experiments were conducted under a purged environment using Nitrogen at a flow rate of 50 ml/min. A baseline measurement was collected using an Al2O3 reference sample of known length. The temperature program for all the measurements was set to start from 50 °C and increase to 1200 °C at a steady rate of 5 °C/min which is the maximum recommended heating rate for better sensitivity of the instrument. Measured results were analysed using the NETZSCH Proteus® software. Samples were prepared by removing the paper liner, then cutting into ≈10 mm length solid rectangular prism samples.

(T ) = (ambient )

(1 +

)

3 dL (T ) L

(2)

3.4.4. Results and discussion – thermo-physical characterization Understanding the thermo-physical properties of fire-resistant grade gypsum plasterboard at both ambient and elevated temperatures is very useful when using them as passive fire protection building materials. Effective specific heat, density and thermal conductivity variations with increasing temperature are the key thermo-physical properties of plasterboards in conjunction with their fire performance. The effective specific heat variation of the three plasterboards was calculated using the DSC and TGA results. The TGA curves were used to obtain the density variations. The density here is the density calculated following the mass loss curve of TGA. The volume change induced with thermal expansion and shrinkage was not considered in calculating the density variation of plasterboards. The density values published for 16 mm fireresistant grade plasterboard by the manufactures were used as the initial density. The manufacturer provided densities were verified by measuring the mass of 10 × 10 × 30 mm3 small rectangular prisms. The calculated densities were within ± 5%. Due to such small deviations from published densities and probable errors of measurements, the manufacturer provided density was used as the initial density of each board. The thermal conductivity variations of the three boards were calculated using the thermal diffusivity data with Equations (1) and (2). The thermo-physical properties of plasterboards are directly related to the chemical composition of the core of plasterboard [7]. The effective specific heat variations of the three gypsum plasterboards are shown in Fig. 4. All the samples exhibited similar effective specific heat versus temperature characteristics with two endothermic peaks at temperatures about 150 °C and 172 °C which are related to two dehydration steps of Gypsum stated earlier in Reactions 1 and 2. The board with the largest gypsum content (Board 3) shows the highest peak values of 14,000 J/kg/°C at 155 °C and 11,500 J/kg/°C at 173 °C. Board 2 with comparatively lower Gypsum content showed its first peak of 10,400 J/kg/°C at 148 °C. Board 2 displays a strong second dehydration peak of 9750 J/kg/°C at 173 °C due to the presence of additional Calcium Sulphate Hemi-hydrate in the form of Bassanite as discussed in

3.4.3. Laser flash analyser (LFA) Thermal diffusivity at elevated temperatures was obtained using laser flash technique from NETZSCH LFA 467 Hyperflash. The instrument measures the thermal diffusivity at pre-programmed temperatures. A heating rate of 20 °C/min and a purged environment using Nitrogen at a rate of 50 ml/min was used for all the measurements. Thermal diffusivity data were obtained at every 50 °C starting from 50 °C and ending at 500 °C 10 mm × 10 mm solid square samples were cut from the core of gypsum plasterboards and were sprayed with a Graphite coating. As gypsum plasterboards are low thermal conductive materials, the sample thicknesses were kept at low values of ≈2 mm. Several samples were needed to test for Board 3 to obtain proper set of thermal diffusivity data due to its highly porous nature (Fig. 1). The thermal conductivity variation with temperature for gypsum plasterboards was calculated using Equation (1), where the temperature dependent thermal conductivity, λ(T) in W/m·K is defined as a function of thermal diffusivity (a(T) in m2/s), specific heat (Cp(T) in J/kg·K), and density (ρ(T) in kg/m3).

(T ) = a (T ) Cp (T ) (T )

mass retention

(1)

Constant specific heat before dehydration of gypsum was used in the calculations. However, density was calculated by considering both mass and volume changes due to heat treatments of samples. The density, ρ(T) was calculated with Equation (2), using obtained mass retention data from TGA and linear thermal expansion data from DIL. Density at temperature T is a function of ambient temperature density ρ(ambient) and temperature dependant thermal expansion, dL/L(T).

Fig. 3. Crystalline structure (a) gypsum (b) vermiculite (c) glass fibres. 4

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Fig. 5. Density and mass retention (a) Board 1 (b) Board 2 (c) Board 3. Fig. 4. Effective specific heat (a) Board 1 (b) Board 2 (c) Board 3.

Table 2 Calculated mass loss percentage of boards for dehydration.

Section 3.3. In the temperature region 380–390 °C, the exothermic peak related to Reaction 3 can be identified. The two small peaks at temperatures 750 °C and 900 °C could be due to the decomposition of other trace minerals such as Aragonite and Sepiolite (Table 1). Usually the mined Gypsum for plasterboard industry can contain impurities with different percentages. As Gypsum is not purified and mined Gypsum is only partially calcined (Reaction 1) in the production line of plasterboards, impurities such as carbonates and silicates can still be present in the final plasterboard product. The gradual increment of effective specific heat curve at temperatures over 1100 °C could be due to impurities and/or Reaction 4 of Gypsum. The initial densities of 16 mm gypsum boards 1, 2 and 3 are 812.5, 787.5 and 784.5 kg/m3, respectively. The apparent density/mass variations with temperature for theses boards are shown in Fig. 5. In Boards 1 and 3 a mass reduction of about 16–17% while a mass reduction of about 14% occurs in Board 2 with two dehydration reactions at temperatures between 115 and 185 °C. The mass reductions due to the dehydration of Gypsum (Reaction 1) and Hemihydrates (Reaction 2) observed can be theoretically calculated using Equation (3) and the results are given in Table 2. Another 4–6% mass reduction occurs after 650 °C following the decompositions of impurities of Gypsum.

Reaction

Board 1

Board 2

Board 3

Reaction 1 Reaction 2 Reactions 1 and 2

13.15 4.5 17.65

10.19 3.86 14.05

13.23 4.47 17.70

Mass loss Reaction 1 = Percentage of Gypsum × 20.9% × 75% Mass loss Reaction 2 = Percentage of Hemihydrates × 20.9% × 25% (3) As seen in Fig. 5, the apparent density/mass variations of these plasterboards are small (total mass loss ≤ 24%) and thus the effects of cracking and shrinkage will be minimised beyond 650 °C. The boards that exhibit rapid reduction in mass are likely to cause premature integrity and insulation failures [22], and can only provide lower FRLs. Thermal expansion curves of the three Gypsum plasterboards from Dilatometer tests are shown in Fig. 6. All three boards showed an initial expansion up to a temperature of about 125 °C. This initial expansion is due to the typical solid material thermal expansion with increasing temperature. The gradual shrinkage in the temperature region of 125–155 °C can match with the reduction in apparent density and/or 5

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Fig. 6. Thermal expansion (a) Board 1 (b) Board 2 (c) Board 3.

Fig. 7. Thermal diffusivity (a) Board 1 (b) Board 2 (c) Board 3.

mass loss of gypsum plasterboards with their dehydration steps (Reactions 1 and 2). The shrinkage starting at about 325 °C is most likely due to the transformation of CaSO4 (III) to CaSO4 (II) (Reaction 3) when compared with effective specific heat (Fig. 4) and apparent density (Fig. 5). Sudden expansion after 400 °C in Board 2 is considered to be due to the expansion of vermiculite and relative behaviour of differently oriented materials surrounding vermiculite [12]. As the last step all three gypsum plasterboards contract continuously, after 800 °C. This is the same temperature region glass fibres lose their mechanical properties. Therefore, this continuous contraction can be identified as the starting point of integrity failure of gypsum plasterboards. As the maximum temperature measurement for LFA 467 is limited to 500 °C, thermal diffusivity data is only available up to that temperature. Fig. 7 shows the measured thermal diffusivity data for the three boards. The thermal conductivity values were then calculated using Equation (1) and the measured thermal diffusivity data. The constant effective specific heat/ambient temperature specific heat for all three plasterboards was taken as 1000 J/kg·K in Equation (1). This value was decided based on recent research studies [15,16] and manufacturer provided room temperature specific heat data. As discussed in Section 3.4.3 the effect of thermal expansion was also considered in the

calculations. Fig. 8 shows the thermal conductivity variations for the three gypsum plasterboards up to 500 °C. The obtained thermal conductivity variation can be divided into four main phases. The explanation is generalized for the three boards and is partially made considering the thermal expansion results. Firstly, it shows a slight reduction up to about 100–120 °C following the solid material expansion region discussed in Fig. 6. This reduction occurs with increase of air pores/void spaces between particles with material expansion. The second phase up to about 170–200 °C also shows a gradual reduction with the mass loss due to dehydration of Gypsum, which again increases the void spaces between particles. In the next phase thermal conductivity remains constant up to about 350–370 °C. In the final phase thermal conductivity gradually increases triggering after material transformation from CaSO4 (III) to CaSO4 (II) (Reaction 3) at the narrow temperature range of about 375–400 °C. This gradual increase of thermal conductivity is found to be due to a material sintering process [16]. Sintering is a process of compacting and forming a solid mass of material by heat without melting it to the point of liquefaction. Sintering leads to increased contact surface between particles, thus the increase in conductivity. Gradual increase of thermal conductivity can be predicted beyond 400 °C until integrity failure of plasterboards 6

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Fig. 9. Idealized effective specific heat for gypsum plasterboards.

4.1. Elevated temperature thermal properties The use of appropriate thermo-physical properties at elevated temperatures is essential in FE heat transfer modelling of LSF wall systems exposed to fire conditions. Specific heat, density and thermal conductivity variation with temperature of wall components together with emissivity and convective coefficients of surfaces are the key inputs for FE heat transfer modelling. To simulate the fire performance of gypsum boards in LSF walls, the measured thermo-physical properties of the three gypsum plasterboards in Section 3.5 were idealized with linear curve fitting to reduce the input data points and thus analysis time. The idealized specific heat, density and thermal conductivity variations for FE modelling are shown in Figs. 9–11. As the samples used for thermal property measurements (≈15 mg in STA and 10 × 10 × 2 mm in LFA) are small without the paper layer and include differences of calculation assumptions, the measured and/or calculated thermal properties for plasterboards are likely to have some deviations from true properties. However, the measured thermal property values can be used to compare the performance of plasterboards alone (used later in Section 5). When simulating the thermal performance of LSF walls exposed to fire, the heat and/or elevated temperature influenced behaviours such as ablation, cracking and fall-off of gypsum plasterboards, migration of moisture vapours, penetration of hot furnace gases into the cavity, etc. are to be considered. However, due to the complexity of modelling these processes, the thermal properties are usually modified to allow for these effects. These elevated temperature influenced behaviours usually alter the heat transfer modes of materials. As thermal conductivity, the only temperature dependant heat transfer mode used in FE heat transfer modelling, the effects due to the above processes are incorporated in idealizing the thermal conductivity.

Fig. 8. Thermal conductivity (a) Board 1 (b) Board 2 (c) Board 3.

considering the constant state of mass loss and thermal expansion variation in Figs. 5 and 6. With integrity failure, plasterboard becomes partially transparent leading to another dominant heat transfer mode, namely heat radiation. These aspects of heat transfer are discussed later under Finite Element (FE) modelling. 4. Numerical modelling Structural and thermal performance evaluations of LSF wall configurations allow fire researchers and designers to determine their FRLs. In recent times, finite element (FE) modelling is increasingly used for this purpose to supplement the results of full-scale fire tests. However, experimental data from full-scale or small-scale fire tests are needed to validate the developed FE models. Fire performance of the three gypsum plasterboards considered in this study and Kolarkar and Mahendran's [17] fire tests of LSF wall systems were simulated using 3D FE heat transfer models developed using Abaqus/CAE Version 6.14-2 [18]. Following the validation using the fire test results in Ref. [17], the FE model was used to acquire the fire performance results of LSF wall systems lined with the three different gypsum plasterboards.

Fig. 10. Idealized density for gypsum plasterboards. 7

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900 °C. The idealized thermal property and temperature values are given in Tables 3–5. In the case of FE heat transfer model validations of LSF wall systems, the thermo-physical properties of an Australian plasterboard (used in Kolakar and Mahendran's [17] fire tests) proposed in Ref. [15] were used. Thermo-physical properties of steel were obtained from Eurocode 3 Part 1–2 [20]. 4.2. Boundary conditions The three heat transfer modes, viz. conduction, convection and radiation, are integrated in FE modelling. The conduction was defined under material properties as thermal conductivity. The effects of convection and radiation for heat transfer were defined by assigning appropriate convective film coefficients (exposed = 25, unexposed = 10 W/m2°C) and emissivity values (exposed, unexposed, cavity = 0.9). The standard fire curve was assigned to the fire exposed side as a boundary condition using an amplitude curve of time-temperature profile of ISO 834 standard fire curve. . The amplitude curve is a measure of change over a time period. The fire exposed side temperature was allowed to follow the amplitude curve by assigning unity to sink temperature. The initial temperature of the models was assigned by defining a pre-defined field for the entire model at ambient temperature, 23 °C. Further, the Stefan-Boltzmann constant of 5.67 × E−8 and absolute zero temperature of −273 °C were assigned to the models.

Fig. 11. Idealized thermal conductivity for gypsum plasterboards. Table 3 Idealized effective specific heat for gypsum plasterboards. Board 1

Board 2

Board 3

Temperature °C

Cp J/(kg·K)

Temperature °C

Cp J/(kg·K)

Temperature °C

Cp J/(kg·K)

23 80 100 113 152 165 173 195 215 250 1200

1000 1270 1510 2280 13,500 10,000 10,800 2000 1210 1000 1000

23 80 100 113 148 160 173 195 215 250 1200

1000 1280 1540 2400 10,400 8600 9750 2000 1220 1000 1000

23 80 100 112 155 168 173 200 225 250 1200

1000 1220 1540 1940 14,000 11,000 11,500 2000 1220 1000 1000

4.3. Thermal model development and validation Previous research studies [15,21] have developed suitable thermal models to simulate the behaviour of load bearing LSF wall systems with single plasterboard lining (1 × 16 mm) and lipped channel section studs of 90 × 40 × 15 × 1.15 mm at 600 mm centres fire tested by Kolakar and Mahendran [17]. However, similar models were developed here and validated using these fire test results to ensure their accuracy before using them in further studies. To reduce the analysis time the heat transfer models of two LSF walls were developed by rescaling the fire tested wall to half of its original height by considering symmetry. This rescaling did not alter the FEA time-temperature curves predicted by full-scale FEA as shown by Rusthi et al. [21]. The model development was undertaken using 8-node linear heat transfer brick elements (DC3D8) and a global mesh density of 20 mm was used with solid sections. Tie constraints were defined at the interface to facilitate the solid to solid heat transfer. Studs and plasterboard sections in contact with studs were meshed into 10 × 20 mm elements to enable better heat transfer between them. The developed 3D FE heat transfer model of single plasterboard lined LSF wall is shown in Fig. 12. Further details of FE heat transfer modelling techniques used can be found in Refs. [15,21]. The measured average time-temperature profiles of the

As the measured data for thermal diffusivity was only available up to 500 °C (due to instrument limitation), the thermal conductivity variations up to 1200 °C were proposed based on available literature [15,16,19]. The effects of mass transfer (moisture movement) differences of fire and ambient sides were incorporated by the sudden drop in idealized thermal conductivity at dehydration temperature of gypsum plasterboards. As described in Section 3.5, thermal conductivity increases gradually up to 350/370 °C as there are no other heat induced chemical reactions. Therefore, a linear thermal conductivity variation was proposed from 350/370 °C–900 °C, by extrapolating the measured thermal conductivity values between 350/370 °C and 500 °C. Further, cracking and ablation induced partial transparency of boards were incorporated by using a sudden increase of thermal conductivity at Table 4 Idealized density for gypsum plasterboards. Board 1

Board 2

Temp °C

Mass Retention (%)

Density kg/m

23 120 155 177 213 655 752 885 935 1200

100 99.7 89.2 84.6 84.1 83.2 81.6 81.0 79.8 78.5

812.5 810.0 725.0 687.0 683.0 676.0 663.0 658.0 648.0 638.0

3

Board 3

Temp °C

Mass Retention (%)

Density kg/m

23 118 155 178 208 672 760 868 960 1100 1200

100 99.7 91.2 86.1 85.6 84.7 82.3 81.5 79.6 79.1 78.0

787.5 785.0 718.0 678.0 674.0 667.0 648.0 642.0 627.0 623.0 614.0

8

3

Temp °C

Mass Retention (%)

Density kg/m3

23 120 163 186 230 680 770 890 960 1100 1200

100 99.4 89.2 83.4 83.4 82.1 80.1 79.0 76.9 76.2 75.2

784.4 780.0 700.0 658.0 654.0 644.0 628.0 620.0 603.0 598.0 590.0

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Table 5 Idealized thermal conductivity for gypsum plasterboards. Board 1

Board 2

Board 3

Temperature °C

Thermal Conductivity W/(m·K)

Temperature °C

Thermal Conductivity W/(m·K)

Temperature °C

Thermal Conductivity W/(m·K)

23 80 150 215 400 900 1000 1200

0.26 0.26 0.12 0.12 0.15 0.25 0.7 1.4

23 80 150 215 350 900 1000 1200

0.25 0.25 0.11 0.11 0.14 0.23 0.7 1.4

23 80 150 225 400 900 1000 1200

0.25 0.25 0.11 0.11 0.15 0.23 0.7 1.4

gypsum plasterboard surface and corresponding FEA results showed a reasonably good agreement as shown in Fig. 13(a). Similarly, the experimental and FEA average time-temperature variations of hot flange (HF), web and cold flange (CF) of studs showed a good agreement as shown in Fig. 13(b).

likely to occur. The ambient side plasterboard surface temperature reaches 200 °C before the cavity side surface reaching integrity failure [2–4]. Therefore FRL of non-load bearing LSF wall systems can be predicted using the plasterboard ambient side time-temperature profiles from FEA, i.e. the time in minutes when the ambient side plasterboard surface reaches the insulation failure limiting temperature of 200 °C. In the case of load bearing walls, time-temperature profiles of both steel stud hot flange and plasterboard ambient side need to be used. In this case, steel stud hot flange reaches its limiting temperature before the ambient side plasterboard surface reaches the insulation limiting temperature [2–4]. Therefore, FRL of load bearing walls is often governed by the time to reach the hot flange limiting temperature of 500 °C in the case of load ratio equal to 0.4. Load ratio is the ratio of applied load on the stud to its ultimate ambient temperature capacity. However, the exact failure time of stud can be obtained from FE structural modelling of stud. The temperature variation profiles of stud from heat transfer modelling of LSF wall is used as input in the structural modelling of stud. To obtain the time-temperature profiles of LSF wall

4.4. FRL of LSF wall systems The validated 3-D FE heat transfer model was used to predict the fire resistant levels (FRL) of load bearing and non-load bearing LSF wall systems lined with the three gypsum plasterboards (Boards 1, 2 and 3) considered in this study. The FRL of non-load bearing walls is defined in terms of integrity and insulation criteria only. The integrity of LSF wall systems is lost with cracking and ablation of plasterboards, which occurs at about 900 °C as stated in Section 4.1. Based on the past fire test results of gypsum plasterboard lined walls [2,17], the chance of an integrity failure of ambient side plasterboard is considered to be minimal while the insulation failure by reaching the maximum limiting temperature of 200 °C (180 °C + room temperature 20 °C) [5] is more

Fig. 12. 3D FE heat transfer model of single plasterboard lined LSF wall (a) mesh details, (b) wall model (c) temperature contours of wall 60 min (d) stud 60 min. 9

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Table 6 FRL of LSF wall systems lined with the three gypsum plasterboards.

1200.0

Temperature (°C)

1000.0 Board 1 Board 2 Board 3

800.0 600.0 400.0

0.0 10

20

Fire_Exp Fire-Cavity_FEA Amb_Exp

30 Time (min)

40

FireISO_FEA Amb-Cavity_Exp Amb_FEA

50

60

Fire-Cavity_Exp Amb-Cavity_FEA

Thermo-physical properties of fire protective boards, namely, specific heat, density and thermal conductivity of the materials, are important parameters affecting the overall system fire performance of LSF walls as evident as evident from fire tests and FE modelling [2,14,17,21]. Taking this into account, this paper proposes a fire performance factor (k-factor) and its variation with temperature as an indicative measure of the fire performance of boards that can be used as a relative classification for fire protective boards. The k-factor defined by Equation (3) is based on thermo-physical properties and is defined as a function of specific volumetric enthalpy (E(T) in J/m3), specific heat (Cp(T) in J/(kg°C)), density (ρ(T) in kg/m3), thermal conductivity (λ(T) in W/m·K) at temperature T and thickness of plasterboard (t in mm) (TA - ambient temperature). Assuming that the above parameters are the most critical ones influencing the fire performance of boards, the kfactor is defined using Equation (3). However, it does not account for the extensive ablation and subsequent integrity failure of plasterboards during fire exposure. As discussed in Section 3.5 the extensive mass loss in the early stages of fire exposure is the main reason for ablation and cracking. Therefore, suitable approximate mass loss limits are also proposed in association with the k-factor to ensure the applicability of Equation (3). It is proposed that the total mass loss by 1200 °C (standard fire temperature) and the mass loss by 200 °C (insulation failure temperature) are limited to 25% and 20% from the initial value, respectively, based on the mass loss of gypsum plasterboards shown in Fig. 5.

Temperature (°C)

1000.0 800.0 600.0 400.0 200.0 0.0 20

30 Time (min)

40

50

60

Fire_Exp

FireISO_FEA

HF_Exp

HF_FEA

Web_Exp

Web_FEA

CF_Exp

CF_FEA

(b) Fig. 13. Comparison of FEA and Experimental time-temperature profiles (a) plasterboard surfaces, (b) hot flange (HF), web and cold flange (CF) of studs.

1200 1000

Temperature (°C)

92.5 89.0 94.0

5.1. Development of standard profile

1200.0

10

57.0 58.0 68.5

5. Fire performance factor (k-factor)

(a)

0

Non-load bearing (200 °C)

system lined with the three gypsum plasterboards considered in this study, the idealized thermal properties proposed in Figs. 9–11 and Tables 3–5 were used as the thermal property inputs. Fig. 14 presents the time-temperature profiles of steel stud hot flange and ambient side plasterboard from these analyses. Table 6 summarizes the FRL of both load bearing and non-load bearing LSF wall systems lined with the three gypsum plasterboards derived from FEA results in Fig. 14.

200.0

0

Load bearing (500 °C)

k (T ) =

E (T ) ×t= (T )

T TA

Cp (T ) (T ) dT (T )

×t

(4)

800 1

Limiting temperature for load

600

0.9

bearing walls – 500 °C

0.8

400 k - factor (× 1011)

Limiting temperature for nonload bearing walls – 200 °C

200 0 0

10

20

30

40 50 Time (min)

60

FireISO

Board1_Amb

Board2_Amb

Board1_HF

Board2_HF

Board3_HF

70

80

90

0.7 0.6 0.5 Board 1

0.4

Board 2

0.3

Board 3

0.2

Board3_Amb

0.1 0 0

Fig. 14. Time-temperature profiles of steel stud hot flange and ambient side plasterboard.

200

400

600 800 Temperature (°C)

1000

1200

Fig. 15. Variation of k-factor with temperature for the three gypsum plasterboards. 10

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Fig. 15 displays the k-factor profiles of the three 16 mm thick gypsum plasterboards calculated using Equation (3) by using the measured thermo-physical properties of each board. However, thermal conductivity was calculated again to develop this fire performance factor. In the previous thermal conductivity calculations, literature/ manufacturer provided ambient temperature specific heat value (1000 J/(kg°C) was used as the constant specific heat value. However, the specific heat obtained at the lowest measureable temperature of 50 °C (1100–1200 J/(kg°C) was used in the calculations. The procedure used in Section 4.1 was used to obtain the measured thermal conductivity variation up to 900 °C. The proposed sudden increase of thermal conductivity at 900 °C for FE modelling to allow for the effects of board cracking was not considered, i.e. thermal conductivity after 900 °C was assumed to be constant until 1200 °C. The mass loss limits were thus proposed to partially account for ablation and cracking. As illustrated in Fig. 15, the k-factor profiles of the three plasterboards show a similar trend with increasing temperature. They follow a sudden increase at 112 °C caused by the enthalpy increase due to Calcium Sulphate Dihydrate (CaSO4·2H2O) dehydration with subsequent reduction in thermal conductivity at the same temperatures. The k-factor profiles of the three gypsum plasterboards from different manufacturers show the same trend as seen in Fig. 15. Determining the “apparent k-factor profile” (Fig. 16) for the standard fire-resistant grade boards was decided by considering the correlation between the calculated k-factor profiles and FRLs summarized in Table 6 for the gypsum plasterboards considered in this study. The predicted time-temperature curves for LSF wall panels in Fig. 14 show that the differences in time-temperature curves and FRLs are small and appear to correlate well with the calculated k-factor profiles in Fig. 15. FEA results show that FRL of Board 2 is slightly lower compared to the other two. This corresponds to the lowest k-factor profile below 525 °C in Fig. 15. After that Board 1 gives the lowest k-factor profile. Therefore, an “apparent k-factor profile” (Fig. 16) was developed using the lowest k-factor at each temperature for the three boards. All three gypsum plasterboard manufacturers claim the same FRLs of 90 and 60 min for non-load bearing and load bearing LSF walls lined with single layer of 16 mm plasterboard, respectively. These claims were proved to be correct by the full-scale fire tests conducted by QUT research group [2,16]. Full-scale fire tests of LSF walls lined with Boards 1 and 2 provided FRLs above 90 min and 60 min for non-load bearing and load bearing LSF walls, respectively [2]. The time-temperature profiles from FE analysis also depict that FRLs of non-load bearing and load bearing LSF walls with all three boards are almost 90 min and 60 min, respectively (Table 6). Further, a parametric study conducted at QUT [14] showed that the FRL of both non-load bearing and load bearing LSF walls is approximately proportional to the plasterboard thickness. Since the proposed apparent k-factor profile is for

2.5

k - factor (× 1011)

2

k - factor (× 1011)

400

600 800 Temperature (°C)

1200

k - factor (× 1011)

FRL – 150 1.5

FRL – 120 FRL – 90

1

FRL – 60 0.5

0 0

200

400

600 800 Temperature (°C)

1000

1200

Fig. 18. Proposed standard k-factor profiles for plasterboards with different FRLs for load bearing LSF walls.

16 mm plasterboard that provides 90 min FRL for non-load bearing LSF walls and 60 min for load bearing LSF walls (Load Ratio = 0.4), the “standard k-factor profiles” for FRLs of 60, 120, 180 and 240 for nonload bearing LSF walls and 60, 90, 120 and 150 were calculated using Equation (4) and are shown in Figs. 17 and 18, respectively. The base FRL in Equation (4) is the FRL of LSF wall system with single 16 mm plasterboard. It is 90 min for non-load bearing and 60 min for load bearing (Load Ratio = 0.4) walls. These standard k-factor profiles can be used as indicators of the overall fire performance of fire-resistant grade plasterboards in LSF wall applications for different FRLs.

Standard k factor profile = Apparent k factor profile ×

FRL Base FRL

(5)

The mass loss limits stated earlier should also be considered when comparing any plasterboard against this proposed standard k-factor profile. Those were, the total mass loss by 1200 °C and the mass loss by 200 °C, to be limited to 25% and 20% from the initial value, respectively. Further, all the thermo-physical property tests of plasterboards should be conducted under the same standard conditions used in measuring the thermo-physical properties in this study. The k-factor profile of any fire-resistant grade plasterboard, calculated using Equation (4), should lie above the appropriate curve proposed in Figure/s 17 and/or 18 for it to be used as lining material for LSF wall systems with proclaimed FRL. If part of the k-factor profile of a given plasterboard lies below the proposed standard in these figures, LSF walls lined with those boards should be tested using the standard

Unsafe Zone

0.3 0.2 0.1 0 1000

1000

2

0.5

600 800 Temperature (°C)

200

2.5

0.6

400

FRL – 60

Fig. 17. Proposed standard k-factor profiles for plasterboards with different FRLs for non-load bearing LSF walls.

0.7

200

FRL – 120

0

Safe Zone

0

1

0

1

0.4

FRL – 180

0.5

0.9 0.8

FRL – 240 1.5

1200

Fig. 16. Apparent k-factor profile for gypsum plasterboards. 11

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composition and thermal property characterization of commonly used gypsum plasterboards from three manufacturers. 3-D FE heat transfer models of LSF wall systems were then developed and validated using available fire test results. The main thermo-physical properties, namely, specific heat, density and thermal conductivity required for FE heat transfer modelling of gypsum plasterboards were proposed based on measured values and used in FE heat transfer analyses. Finally the timetemperature curve variations obtained from these models were used to determine the fire resistant levels (FRLs) of LSF wall systems lined with the three gypsum plasterboards considered in the study. As an overall measure of the fire performance of plasterboards for LSF wall applications, a fire performance factor (k-factor) was proposed as a function of specific volumetric enthalpy, specific heat, density and thermal conductivity at elevated temperatures based on the measured thermal properties and thickness of the board. The “standard k-factor profile” was established based on the correlation between the k-factor and the fire performance of LSF wall panels. The fire performance of LSF wall systems was determined by using the FRL values provided by 3-D FE heat transfer models of LSF walls lined with the three gypsum plasterboards. The k-factor profiles and the time-temperature curves from FE analysis showed a similar behaviour and hence it was concluded that FRL of LSF wall panels lined with the three gypsum plasterboards correlated well with the proposed k-factor profiles. The lowest k-factor values at each temperature for the three gypsum plasterboards was proposed as the standard to predict the fire performance of plasterboards used in LSF wall applications. The standard k-factor profiles were then proposed for non-load bearing and load bearing LSF wall systems with FRLs in the range of 60–240 min. Supplementary conditions on mass loss have also been proposed to allow for the ablation and subsequent integrity failure of plasterboards due to excessive mass loss. Finally, the accuracy of the proposed standard k-factor profiles was verified using two other boards currently available in the market. Further numerical studies and tests are currently under way using other boards with considerable differences in thermal properties to verify the suitability of the proposed standard k-factor profiles. The simplified k-factor approach proposed in this paper assumes that the fire performance of LSF wall systems can be estimated based on the most dominant component of the wall systems, the board lining. Such a single component approach has its limitations but is still useful in assessing the use of unclassified fire protective boards if full scale fire test results are not available.

2.5

k - factor (× 1010)

2

1.5

1

0.5

0 0

200

400

600 800 Temperature (°C) Standard k-factor Profile - 240 Board A Standard k-factor profile - 120 Board B

1000

1200

Fig. 19. Verification of proposed standard profile of overall fire performance of plasterboards.

fire testing procedure given in AS 1530.4 [5] to verify the claimed FRL. The plasterboards with the entire k-factor profile located above the proposed appropriate standard with proper mass loss limits can be considered safe to use in LSF wall applications for fire design with stated FRL. 5.2. Verification of the proposed standard k-factor profile The compliance of the standard k-factor profiles proposed based on thermo-physical properties in determining the fire performance of plasterboards was verified by using two other boards currently available in Australian market. Since no other types of gypsum plasterboards were available, the following two boards were used to demonstrate the applicability of the proposed k-factor approach. The key compounds of the first board is MgO and MgCl2 (Board A – thickness 10 mm) and the second board is made of CaSiO2 (Board B – thickness 20 mm). The manufacturers of these two boards (A and B) claimed 240 min and 120 min of FRL for non-load bearing LSF walls with single layer lining. Hence their k-factor profiles should be above those shown in Fig. 19 for 240 and 120 min FRL. The k-factor profiles of Boards A and B were calculated using the thermo-physical properties given in Rusthi et al. [21] and are also plotted in Fig. 19. LSF walls lined with both Boards A and B were fire tested by our group at QUT [22,23]. Their failure times were 32 min and 137 min, respectively. Based on these fire test results, the k-factor profile of Board A should lie well below the standard profile for FRL 240 min or should fail in mass loss limits. For Board B, k-factor profile should lie above the standard profile for FRL 120 min and should satisfy the mass loss limits. As depicted in Fig. 19, Board B shows an overlap with 120 FRL standard k-factor profile. Therefore, non-load baring LSF wall with Board B can provide FRL of 120 min based on kfactor profiles in Fig. 19, which was also confirmed by the fire test [23]. In the case of Board A, k-factor profile lies below its claimed FRL of 240 min (Fig. 19). Also it failed to satisfy the mass loss limit of 25% at 1200 °C by having 42% mass loss, thus it can be easily decided that Board A will not provide 240 min FRL. This was also confirmed by the fire tests in Ref. [22] (32 min only). These comparisons verify the potential use of the proposed k-factor approach in assessing the fire protective quality of boards and in estimating the fire performance of LSF walls.

Acknowledgements The authors would like to thank Queensland University of Technology (QUT) for providing the necessary research facilities, and Australian Research Council (DP160102879) and QUT for providing the financial support to conduct this research project. The data reported in this paper were obtained at the Central Analytical Research Facility (CARF) operated by the Institute for Future Environments (QUT). Access to CARF is supported by generous funding from the Science and Engineering Faculty. References [1] Australian Building Codes Board, National Construction Code Volume One. Canberra, Australia, (2016). [2] S. Gunalan, P. Kolarkar, M. Mahendran, Experimental study of load bearing coldformed steel wall systems under fire conditions, Thin-Walled Struct. 65 (2013) 72–92. [3] S. Kesawan, M. Mahendran, Fire tests of load-bearing LSF walls made of hollow flange channel sections, J. Constr. Steel Res. 115 (2015) 191–205. [4] A.D. Ariyanayagam, M. Mahendran, Experimental study of load-bearing coldformed steel walls exposed to realistic design fires, Journal of Structural Fire Engineering 5 (4) (2014) 291–330. [5] Standards Australia, AS 1530.4 Methods for Fire Tests on Building Materials, Components and Structures. Sydney, Australia, (2014). [6] International Organization for Standardization, SO 834 -1 Fire Resistance Tests – Elements of Building Construction, Part 1, I, General Requirements, Geneva,

6. Summary of research and findings This paper has presented the details of a study on setting the minimum standards for fire-resistant grade plasterboards used in LSF wall systems. The first phase of this study included the chemical 12

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M.T. Dodangoda, et al. Switzerland, 1999. [7] K.G. Wakili, E. Hugi, Four types of gypsum plaster boards and their thermophysical properties under fire condition, J. Fire Sci. 27 (1) (2009) 27–43. [8] A.H. Buchanan, J.T. Gerlich, Fire Performance of Gypsum Plasterboard, Institution of Professional Engineers New Zealand, Wellington, N.Z, 1997, pp. 160–166. [9] G. Thomas, Thermal properties of gypsum plasterboard at high temperatures, Fire Mater. 26 (1) (2002) 37–45, https://doi.org/10.1002/fam.786. [10] N.B. Singh, B. Middendorf, Calcium sulphate hemihydrate hydration leading to gypsum crystallization, Prog. Cryst. Growth Charact. Mater. 53 (1) (2007) 57–77. [11] S.H. Park, Samuel L. Manzello, Dale P. Bentz, Tensei Mizukami, Determining thermal properties of gypsum board at elevated temperatures, Fire Mater. 34 (5) (2010) 237–250. [12] S.A. Suvorov, V.V. Skurikhin, Vermiculite — a promising material for high-temperature heat insulators, Refract. Ind. Ceram. 44 (3) (2003) 186–193. [13] H. Javangula, Quentin Lineberry, Comparative studies on fire-rated and standard gypsum wallboard, J. Therm. Anal. Calorim. 116 (3) (2014) 1417–1433. [14] ASTM E1269-11, Standard Method for Determining Specific Heat Capacity by Differential Scanning Calorimetry, ASTM International, West Conshohocken, USA, 2005. [15] K. Poologanathan, M. Mahendran, Numerical studies of gypsum plasterboard panels

under standard fire conditions, Fire Saf. J. 53 (2012) 105–119. [16] K. Ghazi Wakili, M. Koebel, T. Glaettli, M. Hofer, Thermal conductivity of gypsum boards beyond dehydration temperature, Fire Mater. 39 (1) (2015) 85–94. [17] P.N. Kolarkar, M. Mahendran, Experimental studies of non-load bearing steel wall systems under fire conditions, Fire Saf. J. 53 (2012) 85–104. [18] Dassault Systems Simulia Corporation, Abaqus/CAE User's Guide. Providence, RI, USA, (2015). [19] J.T. Gerlich, P.C.R. Collier, A.H. Buchanan, Design of light steel-framed walls for fire resistance, Fire Mater. 20 (2) (1996) 79–96. [20] European Committee for Standardization, EN 1993-1-2: 2005, Eurocode 3: Design Of Steel Structures. Part 1-2, General Rules - Structural Fire Design, Brussels, 2005 2005. [21] M. Rusthi, P. Keerthan, M. Mahendran, A. Ariyanayagam, Investigating the fire performance of LSF wall systems using finite element analyses, Journal of Structural Fire Engineering 8 (4) (2016) 354–376. [22] M. Rusthi, A. Ariyanayagam, M. Mahendran, P. Keerthan, Fire tests of Magnesium Oxide board lined light gauge steel frame wall systems, Fire Saf. J. 90 (2017) 15–27. [23] A.D. Ariyanayagam, M. Mahendran, Fire tests of non-load bearing light gauge steel frame walls lined with Calcium silicate boards and gypsum plasterboards, ThinWalled Struct. 115 (2017) 86–99.

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