Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards

Journal of Constructional Steel Research xxx (xxxx) xxx

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Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards Wei Chen a, b, *, Jihong Ye a, Qiang Wang a, Jian Jiang a a b

Jiangsu Key Laboratory of Environmental Impact and Structural Safety in Engineering, China University of Mining & Technology, Xuzhou, 221116, China Xuzhou Key Laboratory for Fire Safety of Engineering Structures, China University of Mining & Technology, Xuzhou, 221116, China

a r t i c l e i n f o

a b s t r a c t

Article history: Received 25 April 2019 Received in revised form 22 October 2019 Accepted 1 November 2019 Available online xxx

Postearthquake fire is one of the most frequent secondary disasters facing mankind. This paper presented a detailed postearthquake fire investigation of cavity-insulated load-bearing cold-formed steel (CFS) walls lined with double-layer gypsum plasterboards on both sides. Five identical full-scale specimens with 3.0 m lengths and 3.0 m widths were constructed. The first two specimens were subjected to axial compression under the condition of ambient temperature or fire, respectively. The other three specimens were first tested under cyclic loading to different drift ratios, and then the fire experiment was carried out. The relationship between the drift ratio and fire resistance time (FRT) is presented, and the following conclusions are drawn. (1) When the drift ratios are less than 2.0%, the reduction of postearthquake FRT is insignificant and should be within 10 min. However, when the drift ratio reaches 3.5%, the residual postearthquake FRT is only 8 min, and the failure mode of wall framing changes to local crushing near the bottom of studs; It is because that local buckling near the bottom of stud occurs from cyclic loading before fire tests, leading to the separation of fire-side sheathing from wall studs. (2) The effect of axial thermal expansion constraints provided by rigid floor slabs to the CFS wall studs is significant under fire conditions. It will accelerate the structural failure of load-bearing CFS walls. (3) For the gypsum-sheathed load-bearing CFS walls, the existence of cavity insulation not only reduces the FRT, but also may change the failure mode of wall studs. © 2019 Elsevier Ltd. All rights reserved.

Keywords: Postearthquake fire experiments Full-scale Cold-formed steel load-bearing wall Earthquake damage New loading condition

1. Introduction The cold-formed steel (CFS) structures have the advantages of high construction efficiency, less steel consumption, good seismic performance and green environment protection, and have been increasingly used in civil engineering. Fig. 1 shows the schematic diagram of a typical CFS structure. In China, such CFS structures are allowed to be constructed not exceeding 24 m in height and 6 stories in number. Because the thickness of CFS framing is usually only 1e2 mm, the heat transfer of CFS is fast and the fire resistance investigation becomes very important for CFS structures. Meanwhile, multihazard interaction investigation is an important issue for disaster reduction in civil engineering and postearthquake fires are one of the most frequent secondary disasters facing mankind. Recently, Roy et al. conducted a collapse experimental investigation

* Corresponding author. Jiangsu Key Laboratory of Environmental Impact and Structural Safety in Engineering, China University of Mining & Technology, Xuzhou, 221116, China. E-mail address: [email protected] (W. Chen).

of a single-story CFS building under severe fires [1]. It was found that the CFS roof collapsed inward asymmetrically at a time of only 21.5 min of fire exposure and the gypsum-sheathing CFS walls did not collapse in the end of fire test. Besides, the university of California, San Diego (UCSD) conducted an earthquake and postearthquake fire experiment of a large-scale six-story CFS building using a large outdoor shaking table and a pool fire in 2016 [2]. The experimental results were helpful to evaluate the earthquake performance of mid-rise CFS structures, as well as the earthquakedamaged CFS components’ response to fire. However, due to the high cost and long period, the experimental investigation of whole CFS buildings is very limited. Most studies on earthquake and fire performance of CFS structures focused on their main components [3e21]. In CFS structures, the CFS walls, which consist of CFS framing lined with sheathing boards on both sides, are the main loadbearing and lateral resistance components, and play an important role in the mechanical performance of the whole structure. Fig. 2 shows the schematic diagram of a typical CFS wall. In Fig. 2, the CFS framing is built by assemble the lipped or unlipped channel section studs with the top and bottom tracks made of lipped or

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walls and are very common in mid-rise CFS structures. A detailed postearthquake fire experiment of such wall assemblies is performed. Some quantitative earthquake damage is determined through the drift ratio, which is obtained from the cyclic loading tests. A new loading condition is developed that can represent the actual situation in mid-rise CFS structures. Subsequently, the influences of earthquake damage, loading conditions and cavity insulation on the fire performance of such wall assemblies are discussed in detail.

2. Test program 2.1. Test devices & improved loading condition

Fig. 1. Schematic diagram of a typical CFS structure.

Fig. 2. Cavity-insulated cold-formed steel wall lined with double layers of sheathing on both sides.

unlipped channel sections, using self-drilling wafer head screws. Meanwhile, the sheathing boards, including gypsum plasterboard (GP board), calcium-silicate board, oriented strand board and steel sheet et al. are usually fixed to the CFS framing by self-drilling bugle head screws. Among these sheathing boards, GP board is the most widely used, because of the low price, good fire and decorative performance. In general, it is necessary to adopt double layers of sheathing on both sides of the load-bearing CFS frame to satisfy the fire resistance ratings of mid-rise buildings. Moreover, to meet the requirements of thermal and sound insulation, some filling materials, such as rock wool, glass wool and aluminum silicate wool, are usually used as cavity insulation in CFS walls. At present, various cyclic loading tests and fire experiments of CFS walls have been conducted to determine the earthquake [3e10] and fire performance of such assemblies [11e20]. However, the investigations of the interaction behavior of CFS walls during a combined earthquake and fire are still very limited. The National Institute of Standards and Technology (NIST) performed full-scale experiments of steel-sheathed shear walls subjected to seismic and fire loads, and investigated the influence of fire loads on the lateral resistance of such wall assemblies [21]. It is still blank for the investigation of failure mechanism and fire resistance limit of earthquake-damaged CFS load-bearing walls under postearthquake fire conditions. The subject of the present investigation is the cavity-insulated load-bearing CFS walls lined with double layers of GP boards on both sides. Such walls can be used as the load-bearing internal

Fig. 3 showed a conventional loading condition for fire experiments of CFS walls [11,12]. It involved using a moving distributive beam to transmit the vertical load to the wall studs. Moreover, the vertical load was kept constant during the fire experiments. However, in CFS structures, each wall stud is separately subjected to the vertical load. Therefore, the conventional loading condition cannot obtain the actual load for each wall stud. In addition, steel deck floor slabs are very common in mid-rise CFS structures and have much higher flexural rigidity than the traditional OSB floors in lowrise CFS structures. During the fire exposure of a CFS wall, the axial thermal expansion load provided by the rigid floor to the wall studs might become nonnegligible. The conventional loading condition cannot consider this axial thermal expansion load. In addition, it should be noted that the current Chinese code requires consideration of the thermal expansion load in the fire design of steel members [22]. Hence, an improved loading condition was proposed that could represent the actual situation in the fire experiments of CFS walls. Fig. 4 shows the test devices and improved loading conditions. In Fig. 4b, a multi-jack synchronous loading system that was specially designed for the present fire experiments, which enabled each jack to apply independent loads. Using this synchronous loading system, the axial displacement of each jack remained constant during the initial stage of fire exposure to account for the axial thermal expansion constraints on the wall studs; during the final stage of fire exposure, the synchronous loading system will be switched to the force control mode when the measured axial load of the wall stud rapidly decreases and approaches the initial value. In this way,

Fig. 3. Conventional loading conditions from our previous fire experiments of loadbearing CFS walls [11,12].

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 4. Test devices and improved loading conditions.

the axial thermal expansion constraints provided by the rigid floors would be considered throughout the fire experiments. In addition, a moving steel beam (Fig. 4a) was specially designed and separately attached to each jack and corresponding wall stud. In contrast to the moving distributive beam (Fig. 3), the present moving beam did not transmit the vertical load and only provided lateral constraints to the top of the wall specimen. Hence, the loading system and moving beam directly applied the vertical load to each wall stud. In the present experiment, the fire environment is provided by a large-scale vertical furnace system with internal spaces that are 3.0 m in length, 3.0 m in width and 1.5 m in depth, as shown in Fig. 5. The vertical furnace system has six gas burners, eight S-type thermocouples (maximum working temperature of 1600  C) and two high-temperature video cameras. The furnace also has an electric open-close mechanism and can move along the iron rail on the ground. In addition, the vertical load of the wall specimen is

Fig. 5. Large-scale vertical furnace system at the China University of Mining & Technology.

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applied by six oil-jacks and the synchronized loading system, as shown in Fig. 4b and c. The horizontal cyclic loading is applied by the electrohydraulic servo actuator, as shown in Fig. 4d. 2.2. Specimen design Five identical full-scale CFS wall specimens (W1eW5) with 3.0 m lengths and 3.0 m widths were constructed for the present experiments. Fig. 6 shows the configuration of the wall section. The steel frame of each specimen was built by using CFS lipped channel studs (C89 89 mm  40 mm  12 mm  1.0 mm) with a nominal yield strength of 550 MPa. Built-up back-to-back lipped channels were used for the chord studs, and a single lipped channel was used for the middle studs and tracks. The steel studs were assembled with the top and bottom tracks by using 16 mm long self-drilling wafer head screws. Moreover, 100-mm-thick rock wool (density of 60 kg/m3) was filled into the cavity of each specimen. Fig. 6 also shows the details of the hold-down, which has a thickness of 5 mm. Each chord stud had two hold-downs to minimize the uplift deformation of the wall assemblies during cyclic loading. Fig. 7 shows the construction details of wall specimens. Each specimen had double layers of fire-resistant gypsum plaster (GP) boards on both sides and the dimensions of the GP board were 3000 mm  1200 mm  12 mm. Considering the convenience of construction, the GP boards were vertically attached on the CFS frame (Fig. 7b and c). In order to improve the fire resistance performance, a staggered arrangement of GP board joints was used on both sides of the steel frame. In addition, all the face layer GP boards were fixed to the frame and the base layer boards by 45 mm long self-drilling bugle head screws. The screw spacing between the face-layer sheathing and steel frame was 150 mm, and the corresponding screw edge distance was 10 mm for the middle studs and 20 mm for the chord studs and tracks.

2.3. Assembly of the wall specimen To facilitate on-site assembly, the screw holes on the CFS frame were set in the factory in accordance with the construction details (Fig. 7) of each specimen, as shown in Fig. 8a. After the assembly of the steel frame, the K-type thermocouples were welded to the steel stud by using a hot spot thermocouple welder, as shown in Fig. 8b. Subsequently, the steel frame was mounted into the loading frame by bolting the top and bottom tracks to the bottom supporting beam and top moving beam, respectively, as shown in Fig. 8c. Then, the GP board and rock wool insulation were constructed (Fig. 8d and e). After the construction of the wall specimen and the arrangement of sensors, the axial compression test or fire test can be conducted. If the specimen is first subjected to cyclic loading, the left and right column of the loading frame must be temporarily removed, and the top moving beam must be connected to the actuator. After the cyclic loading test, the left and right columns of the loading frame were reinstalled, and the fire test was conducted. Fig. 9 shows the state of preparation for the present experiments.

2.4. Temperature and deformation monitoring For the present axial compression test and fire tests, 8 linear variable displacement transducers (LVDTs) were arranged along the height of each specimen to monitor the out-of-plane deformation, as shown in Fig. 10a; the vertical deflection of each wall stud was recorded by an additional 6 LVDTs, as shown in Fig. 4c. Moreover, in the fire tests, a total of 114 K-type thermocouples were used to determine the temperature profile of each specimen, as shown in Fig. 10b. In addition, 4 laser displacement transducers were used to record the in-plane deformation of each specimen, as shown in Fig. 10c.

Fig. 6. Configuration of the present CFS wall specimen.

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Fig. 7. Construction details of the present wall specimen.

2.5. Test procedure

continued until the structural failure of specimen W1.

In the present five full-scale CFS wall specimens, specimen W1 is used for the axial compression test to determine the corresponding axial load-bearing capacity at ambient temperature; specimen W2 is used for the fire experiments to determine the fire resistance performance of such load-bearing walls without earthquake damage, and the other three specimens (W3, W4 and W5) are used for postearthquake fire experiments, in which the cyclic loading test is performed before the fire experiment to investigate the influence of earthquake damage on the fire performance of such load-bearing walls. Table 1 shows an overview of the present experiments. In the cyclic loading tests, the drift ratio is used to provide quantitative earthquake damage. Specimen W5 is subjected to a drift ratio of 3.5%, which represents the limit state of collapse for CFS structures subjected to earthquakes [2,21]. Specimens W3 and W4 are subjected to drift ratios of 1.0% and 2.0%, which represent the states before and after the ultimate shear strength of such walls during an earthquake, respectively. Hence, there are three types of test procedures for the present wall specimens, as described below.

2.5.2. Cyclic loading tests A displacement-controlled loading mode was adopted for specimens W3, W4 and W5. The CUREE basic loading protocol in ASTM E2126-11 was used for the present cyclic loading tests [23]. The loading procedure involved displacement cycles grouped in phases with incrementally increasing displacement levels. The horizontal loading rate was 1.5 mm/s, and the detailed loading description is shown in Table 2 and Fig. 11. 2.5.3. Fire tests Specimens W2, W3, W4 and W5 were exposed to ISO 834 fire from one side. First, an axial compression load was gradually applied to all the studs (15 kN for each stud). The load was kept constant at ambient temperature for approximately 15 min. Subsequently, the furnace was started and followed the ISO 834 standard time-temperature curve. The fire resistance time of each wall specimen was determined based on the criteria of structural failure, insulation and integrity failure [24]. 3. Test results

2.5.1. Axial compression test at ambient temperature A force-controlled loading mode was adopted for specimen W1. The vertical load was synchronously applied to all six wall studs of W1. Each step of the vertical load was approximately 2.0 kN. After each application step of the vertical load, the total vertical local was kept constant for approximately 2 min. Subsequently, the next step of vertical load was applied. The force-controlled loading mode was

3.1. Axial compression test at ambient temperature (W1) An axial compression test was conducted on specimen W1 at ambient temperature. During the loading process, the maximum out-of-plane deformation was only 2.5 mm at the mid-height of the specimen. When the load reached 55.41 kN, the No. 3 middle stud

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 8. Assembly of steel frame.

(Stud3 in Fig. 10a) was the first to lose the vertical load-bearing capacity, and its corresponding jack exited the working state. Subsequently, studs No. 2, 4 and 5 lost their bearing capacities one after the other. When loaded to 108.43 kN, the No. 1 chord stud (Fig. 10a) lost its bearing capacity. Later, the No. 6 chord stud was also damaged, and the test was stopped. Fig. 12 shows the loadvertical displacement curve of studs No. 3 and No. 1. During the test, local crushing of the face-layer GP board was identified near the loading (top) end of the wall studs (Fig. 13a). After removing the sheathing boards for visual observation, local bucking was found on

the top of the stud section (Fig. 13b). Some studs also exhibited local bucking at the bottom of the wall studs (Fig. 13c). In addition, there was no obvious buckling phenomenon at the other heights of the wall studs. 3.2. Cyclic loading tests (W3, W4 and W5) Before the fire experiments, specimens W3, W4 and W5 were first subjected to cyclic loading tests, in which the quantitative earthquake damage on the wall specimens was given through

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Fig. 9. Preparation of present experiments.

different drift ratios. One of the most important works in the present investigation is to evaluate the impact of these quantitative earthquake damages on the fire resistance performance of the wall. Under cyclic loading, the damage to the present wall specimens was mainly observed as the damaged GP board near the corner of the specimen, opening vertical joints of the sheathing board and sheathing board screws pulled through the board. These damages were introduced as follows: 3.2.1. Damaged GP board near the corner of the specimen When the drift ratio of the specimen reached 1%, the sheathing board near the corner of the specimen started to generate diagonal cracks (Fig. 14a). When the drift ratio of the specimen reached 2e2.25%, the sheathing board near the corner of the specimen broke and partially fell off (Fig. 14b). Thereafter, notable relative sliding occurred between the sheathing board and chord stud during the cyclic loading process (Fig. 14c). 3.2.2. Opening vertical joints of the sheathing board As shown in Fig. 7, the present wall specimens displayed vertical joints, and there was no obvious gap in the vertical joints of wall specimens before testing (Fig. 15a). Generally, the gap in the sheathing board joints of the load-bearing CFS wall would increase as the drift ratio increased. During cyclic loading, when the drift ratio of the specimen reached 1%, the opening of the face-layer sheathing board joints was approximately 0.8 mm (Fig. 15b). When the drift ratio reached 2%, the opening of the face-layer sheathing board joints was approximately 1.6 mm (Fig. 15c). When the drift ratio reached 3.5%, due to unexpected presence of paper tape (which shall be performed after completing the cyclic loading test) on the face-layer sheathing board joints of specimen W5 before testing, an accurate observation of the opening of the sheathing board joints cannot be made (Fig. 15d), but it can be determined that the actual openings of the board joints are less than 3 mm. In addition, it is not difficult to find that the opening of the vertical joints in the present wall specimens is substantially less than that identified by other existing tests of similar wall types [3e8]. This difference is mainly attributed to the fact that the existing similar tests mostly adopted the configuration of singlelayer sheathing on the same side of the steel frame, while the present specimens adopted the configuration of double-layer

sheathing on the same side of the steel frame from the perspective of fire-resistant performance. Based on our previous shearing tests of CFS screw connections with GP sheathing [25], the shear strength and shear deformation capacity of screw connections with double-layer GP sheathing were much higher than those with only single-layer GP sheathing, as shown in Fig. 16. The red curve in Fig. 16 represents the shear force-displacement curve of the screw connections between double-layer 12 mm fire-resistant GP sheathing and a 1.0 mm G550 CFS sheet at ambient temperature, whereas the black curve represents the force-displacement response of the screw connections between single-layer 12 mm fire-resistant GP sheathing and a 1.0 mm G550 CFS sheet. 3.2.3. Screws pulled through the sheathing board When the drift ratios of the specimens reached 2%, the connecting screws along the top and bottom tracks pulled through the GP board (Fig. 17a), implying a separation between the face-layer sheathing boards and the top & bottom tracks of the wall specimens. Thereafter, under cyclic loading, rotational deformation of the sheathing board became visible. When the drift ratios of the specimens reached 2.5%, four rows of screws from the bottom of stud, which was approximately 400 mm from the stud bottom, pulled through the board (Fig. 17b and c). When the drift ratios of the specimens reached 3%, there was no obvious change regarding the screw connections along the middle stud, and the screws on the chord stud, which was approximately 550 mm from the stud bottom, pulled through the board (Fig. 17d). In particular, when the drift ratios of the specimens reached 2.5%, four rows of screws near the stud bottom pulled through the board, which actually reflects two important experimental phenomena. (a) Local buckling already occurs on the middle stud approximately 400 mm from the stud bottom (where the screws pulled through the board). As the chord stud is restrained by the hold-down, the local buckling of the chord stud occurs approximately 550 mm above the stud bottom (Fig. 17e). (b) The sheathing board separates from the steel frame on the bottom of the stud, and an obvious gap is identified (Fig. 17e). The above two important experimental phenomena will substantially impact the postearthquake fire resistance of the present wall specimens. The detailed discussion is given in the next section. Fig. 18 shows the hysteretic curves of specimens W3, W4 and

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 10. Arrangement of the displacement sensors and thermocouples in the experiments.

W5 under cyclic loading. According to the curves, the average shear ultimate load is 75.20 kN (25.07 kN/m) for the load-bearing CFS wall with two layers of GP boards on the same side, and the

corresponding average drift ratio is approximately 1.59% when the shear ultimate load is given. Taking the previous experimental studies as a reference, the shear ultimate load is approximately

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Fig. 10. (continued).

Table 1 Overview of the present experiments of the CFS wall specimens. No.

Test Step

Description

Axial load/fire curve/cyclic loading rate

W1 W2 W3

Step Step Step Step Step Step Step Step

Axial loading at ambient temperature Fire experiments with axial loading Cycling to 1% drift ratio Fire experiments with axial loading Cycling to 2% drift ratio Fire experiments with axial loading Cycling to 3.5% drift ratio Fire experiments with axial loading

e 15 / 15 / 15 / 15

W4 W5

1 1 1 2 1 2 1 2

kN per stud/ISO /1.5 mm/s kN per stud/ISO /1.5 mm/s kN per stud/ISO /1.5 mm/s kN per stud/ISO

834 curve/d 834 curve/d 834 curve/d 834 curve/d

Note: ‘d’ represents null.

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Table 2 Amplitudes and steps of the cyclic loading test. Step

Number of primary cycle/trailing cycle

Amplitude of primary cycle/mm

Amplitude of trailing cycle/mm

Drift ratio/%

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18

6/d 1/6 1/6 1/3 1/3 1/2 1/2 1/2 1/2 1/2 1/2 1/2 1/2 1/2 1/2 1/2 1/2 1/2

1.5 2.25 3 6 9 12 21 30 37.5 45 52.5 60 67.5 75 82.5 90 97.5 105

e 1.68 2.25 4.5 6.9 9 15.9 22.5 28.2 33.9 39.3 45 50.7 56.4 61.8 67.5 73.2 78.9

0.05 0.075 0.1 0.2 0.3 0.4 0.7 1 1.25 1.5 1.75 2 2.25 2.5 2.75 3 3.25 3.5

Note: ‘d’ represents null.

10.68e15.29 kN/m for a CFS wall with a single-layer GP board [3,5]. Hence, the shear capacity of the double-layer GP sheathed CFS wall is approximately twice that of the single-layer GP sheathed wall. However, an obvious no-load slip can still be observed from the hysteretic curves of the present specimens, which is mainly attributed to the expanded screw holes between the sheathing board and steel frame. Therefore, the hysteretic energy-dissipating capacity of the GP-sheathed CFS wall is not desirable.

3.3. Fire experiments of load-bearing wall specimens (specimens W2, W3, W4 and W5)

Fig. 11. Displacement loading in the cyclic loading test.

Fig. 12. Axial load-vertical displacement curves of specimen W1.

3.3.1. Visual observation A fire experiment was conducted with specimen W2, and postearthquake fire experiments were conducted with specimens W3, W4 and W5. Specimens W2, W3, W4 and W5 displayed structural failure under ISO 834 fire exposure without losing insulation or integrity, and their corresponding fire resistance times were 48 min, 41 min, 40 min and 8 min, respectively. The fire experimental phenomena for specimens W2, W3 and W4 were similar: intermittent white smoke emitted from the ambient side of the specimens after 2e3 min of fire exposure. During the final stage of the fire experiments, visible out-of-plane deformation towards the ambient side was gradually displayed on the specimens. Figs. 19e21 show the postfire visual observations of specimens W2, W3 and W4, among which the following phenomena were noted: ① On the fire side, the joints clearly opened and the GP boards partially fell off (Figs. 19a, 20a and 21a). However, the fall off may also be caused by cooling after the fire experiment, which shall be determined based on the analysis of the timetemperature curve of the specimens in the following section. ② The rock wool cavity insulation partially blackened (Figs. 19b, 20b and 21b), implying a higher temperature in the corresponding region. ③ Local buckling of hot flanges occurred approximately 350e700 mm from the top of the stud (Figs. 19d, 20d and 21d), and flexural deflection of steel studs towards the ambient side was identified (Figs. 19c, 20c and 21c). In addition, some specimens displayed local buckling at the bottom of the stud (Fig. 21e) and fracture of the top tracks, which were most likely due to uneven vertical displacement among the studs (Figs. 20d and 21f).

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Fig. 13. Visual observation of specimen W1.

Fig. 14. Failure phenomenon on the corner of the face-layer GP board corresponding to different drift ratios.

Fig. 16. Load-displacement curves of CFS screw connections with single-layer and double-layer gypsum sheathing [25].

Fig. 15. Opening GP board joints of wall specimens corresponding to different drift ratios.

In contrast to specimens W2, W3 and W4, specimen W5 lost its bearing capacity after only 8 min of fire exposure. Obvious expansion occurred at the gap between the sheathing board and steel frame on the bottom of specimen W5 (Fig. 22a). The integrity of the B1-layer and B2-layer GP boards on the fire side was well maintained, i.e., these boards did not fall off. Local crushing on the stud section was observed on the stud 200e400 mm from its bottom (Fig. 22c), and the steel frame showed a small amount of flexural

deflection towards the ambient side (Fig. 22b). In particular, the destruction location of the stud was the same as the local buckling position of the steel frame from the cyclic loading test. In addition, there was no buckling phenomenon on the top or bottom of the studs.

3.3.2. Time-temperature curves Fig. 23a, b and 23c show the time-temperature curves of specimens W2, W3 and W4 500 mm from the top (near the buckling position of the steel frame). Fig. 23d shows the time-temperature curve of specimen W5 500 mm from the bottom. The following definitions apply to Fig. 21: ISO 834 represents the ISO 834 standard time-temperature curve; “FS” represents the average temperature

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 17. Screw heads pulled through the GP board under different drift ratios.

(thermocouples T1, T31, T49, T73 and T91) on the fire side of the B1layer sheathing; “B1eB2” represents the average temperature (thermocouples T2, T32, T50, T74 and T92) on the interface of the B1 and B2 layers of sheathing; “B2eCI” represents the average temperature (thermocouples T33 and T75) on the cavity insulation of the B2-layer sheathing; “CIeB3” represents the average temperature (thermocouples T34 and T76) on the cavity insulation of the B3-layer sheathing; “B3eB4” represents the average temperature (thermocouples T8, T35, T56, T77 and T98) on the interface of the B3 and B4 layers of sheathing; “Amb” represents the average temperature (thermocouples T36, T78, T30, T42, T84, T48 and T90) on the ambient surface of the B4-layer sheathing; and “HF”, “MW” and “CF” represent the average temperature (thermocouples T3, T4, T51, T52, T93 and T94) on the hot flange of the steel frame, the average temperature (thermocouples T5, T53 and T95) on the stud web and the average temperature (thermocouples T6, T7, T54, T55, T96 and T97) on the cold flange of the steel frame, respectively. Fig. 24 shows the temperature distribution along the height of No. 4 middle stud of the four specimens, in which the following definitions apply: “HF-top” represents the average temperature (thermocouples T51 and T52) on the hot flange 500 mm from the top; “HF-mid” represents the average temperature (thermocouple T59 and T60) on the hot flange 1500 mm from the bottom; and “CF-bot” represents the average temperature (thermocouples T70 and T71) on the cold flange 500 mm from the bottom. The detailed features of the time-temperature curve for GPsheathed CFS walls under ISO 834 fire conditions have been given in many studies; thus, these features will not be illustrated here due to limited context. However, the following aspects need to be noted

in Figs. 23 and 24: (1) In Fig. 23a, for specimen W2, during the early stage of fire exposure, the temperature on the hot flange of the steel frame (HF) is very close to that on the B2eCI interface. Later, the HF temperature gradually becomes higher than the B2eCI interface temperature, which is probably due to the following reasons: ① The opening vertical joints of the fireside GP board at the hot flange of the steel frame accelerate the temperature increase on the hot flange, whereas there are no vertical joints on the B2eCI interface; ② The coldformed steel and rock wool insulation material have different thermophysical properties at elevated temperatures. (2) If there are substantial temperature gradients between the time-temperature profiles of the specimen and the furnace temperature, it can be determined that the sheathing board and insulation material at the corresponding position do not fall off throughout the fire experiment. Taking specimen W4 as an example, according to Fig. 23c, during the final stage of the fire experiment, the temperature on the B1eB2 interface (thermocouples T2, T38, T74 and T80) and the HF temperature are substantially lower than the temperature inside the furnace. Similarly, it can be concluded that the fire-side sheathing boards of the present four specimens (W2~W5) do not fall off throughout the experiment. Therefore, the partially separated fire-side GP boards in Figs. 19e21 are probably caused by the cooling shrinkage of the GP board after the fire experiment.

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 18. Hysteretic curves of specimens W3~W5 under cyclic loading.

(3) In Fig. 24a, b and 24c, the top and mid-height temperatures of the steel frame stud (e.g., HF-top and HF-mid) are generally higher than the temperature of the stud bottom (HF-bot). This temperature difference is due to the upward movement of hot air flow in the furnace, which leads to the temperature at the mid-height and above in the furnace being slightly higher than that of the furnace bottom. As a result, more obvious cracks are displayed on the fire-side GP board at the mid-height and above in the specimen, which results in a faster temperature increase at the corresponding height of the steel frame stud. (4) In Fig. 24d, for specimen W5, after 3 min of fire exposure, the temperature on the hot flange 500 mm from the bottom of the stud is substantially higher than that on the B1eB2 interface. After 8 min of fire exposure, the temperature on the mentioned hot flange is close to 550  C. This finding indicates that the heat in the furnace can be directly transferred to the hot flange at the bottom of the steel frame stud without passing the fire-side GP board. This phenomenon is caused by the separation between the fire-side sheathing board and the steel frame at the bottom of the wall assembly under cyclic loading (Fig. 17e). Therefore, the earthquake damage has a direct impact on the fire resistance time of specimen W5. Hence, based on the above visual observations and the timetemperature curves of the specimens (Figs. 19e24), the failure phenomena of the fire experiments for the four specimens are

summarized in Table 3. In this table, Lhot-F represents the local buckling of the hot flange and flexural failure of the steel frame towards the ambient side, and LC represents the local crushing on the stud section. The failure positions of the wall stud all refer to the distance from the stud top to the failure points. 3.3.3. Time-dependent load-displacement curves Fig. 25 shows the time-axial load curves of the middle stud for all specimens. In this figure, the red horizontal line represents the initial axial load, which is 15 kN per stud. The time-dependent axial expansion load of the middle stud can be determined by subtracting the initial axial load from the measured axial load. The figure shows that the studs of all specimens bear substantial axial thermal expansion load during the experiment, and the axial thermal expansion load gradually increases as the temperature of the steel frame increases during the early stage of the fire experiment. Taking specimen W2 as an example, after 33 min of fire exposure, its axial thermal expansion load reaches 10.6 kN, which is close to the initial axial load of the stud. During the final stage of the fire experiment, as the temperature of the steel frame stud continuously increases, its material properties quickly degenerate at elevated temperatures, and the axial thermal expansion load that the stud can bear also gradually decreases until the measured axial load of the stud becomes lower than the initial load; thus, the wall loses its bearing capacity. Fig. 26 shows the time-dependent out-of-plane deflection of the No. 4 middle stud for all specimens. In this figure, Stud4-top represents the out-of-plane deflection of the No. 4 stud 500 mm from

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 19. Postfire visual observation of specimen W2.

the top, and Stud4-mid represents the out-of-plane deflection of the No. 4 stud at mid-height (1500 mm from the top). The negative out-of-plane deflection represents the deformation of the specimen towards the ambient side, and the positive out-of-plane deflection represents the deformation of the specimen towards the fire side. Based on the figure, the following observations can be noted: During the early stage of the fire experiment, specimens W2, W3 and W4 displayed out-of-plane deformation towards the furnace due to the temperature gradients on the stud section. During the final stage of the fire experiment, local bulking occurred on the hot flange of the steel stud approximately 200e700 mm from the top, which resulted in the out-of-plane deformation of the specimen towards the ambient side until the bearing capacity was lost. When specimens W2, W3 and W4 lost their bearing capacities, the out-of-plane deformations of the specimens 500 mm from the stud top were 27 mm, 24 mm and 66 mm, respectively. Specimen W5 was only exposed to fire for 8 min, and there was no obvious out-of-plane deformation at the mid-height of its stud, and the maximum out-of-plane deformation of the stud 500 mm from the bottom was 4 mm.

4. Discussion 4.1. Impact of earthquake damage on the fire resistance time of the wall Fig. 27 shows the relationship between the drift ratio and fire resistance time of the present GP-sheathed CFS walls based on the postearthquake fire experiments. Although the present investigation does not perform a postearthquake fire experiment under a drift ratio of 2.5%, considering that the earthquake damage of the CFS wall under a drift ratio of 2.5% is similar to that under a drift ratio of 3.5%, the residual fire resistance time of the CFS wall after experiencing a drift ratio of 2.5% is also adopted as 8 min from a conservative perspective. The following observations can be made from Fig. 27: (1) When the drift ratio given on the wall specimens (W3 and W4) is within 2%, the minimum fire resistance time is only 8 min less than that of the wall specimen (W2) without earthquake damage, implying that there is no obvious impact from the earthquake damage (i.e., opening vertical joints of sheathing board, cracked sheathing board near the

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 20. Postfire visual observation of specimen W3.

corner of the wall, screws along the tracks pull through the board) under a drift ratio of 2.0% on the fire resistance time of the wall. The reasons for these findings are as follows: ① The cracked sheathing board near the corner of the wall and the pulled through screws along the tracks are local damages that have no obvious impact on the temperature distribution of the CFS wall; ② The openings of the vertical board joints of CFS walls are rather small under cyclic loading. Fig. 28 shows the states of vertical board joints for specimen W2 without earthquake damage in the early stage of the fire experiment, which were obtained by the high-temperature camera. The burning of the paper tape along the fire-side vertical board joints occurred after the fire experiment started; in addition, after approximately 10 min of fire exposure, the vertical joints of the fire-

side GP board substantially opened and became larger than the diameter (8.9 mm) of the screw head, implying that the opening of the vertical joints on the fire-side GP board are inevitable phenomena for a GP-sheathed CFS wall during the early stage of fire exposure. In addition, the two layers of sheathing boards on the same side of the CFS wall assembly are arranged with staggered joints, which further reduces the impact of opening vertical joints on the fire resistance of the CFS walls. Therefore, the timetemperature curves of the hot flange are very similar for specimens W2, W3 and W4 (Fig. 29). Considering that the openings in the vertical board joints are less than 3 mm under the drift ratio of 3.5%, the following wall damages (including opening vertical joints of sheathing board, cracked sheathing board near the corner of the wall, pulled through screws on the tracks and sheathing board stuck into the boards) of the load-bearing CFS wall (covered with double-layer GP boards and lined with insulation) under cyclic

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 21. Postfire visual observation of specimen W4.

loading do not have an obvious impact on the fire resistance of the present wall, and the reduction in the fire resistance time should be within 10 min. (2) After experiencing a drift ratio of 3.5%, the residual FRT of specimen W5 is only 8 min, which is substantially lower than that of specimen W2 without earthquake damage. The reasons are as follows: ① After the cyclic loading test, an obvious gap is generated between the bottom sheathing board on the fire-facing side of assembly W5 and the steel frame (Fig. 17e); hence, the fire-side GP board cannot provide fire protection for the bottom area of the steel stud during the fire test. As a result, the heat in the furnace can be directly transmitted to the hot flange of the stud through the gap between the sheathing board and steel frame (Fig. 30), which leads to a sharp temperature increase at the bottom of the stud (Fig. 24d) and a rapid decrease in the material properties of the CFS; ② Local buckling occurs on the middle stud at a point near the stud bottom under cyclic loading (Fig. 17e); thus, the axial compression loading capacity of the stud dramatically decreases after the cyclic loading test. Hence, the residual FRT of specimen W5 became less than a naked nondestructive steel frame under fire conditions. Hence, for the design practice of CFS structures, the story drift ratios of CFS structures are usually less than 2.0% under the design earthquake actions. At this time, the degeneration of fire performance of cavity-insulated load-bearing CFS walls lined with

double-layer GP sheathings on either side is insignificant after the design earthquake actions and the reduction in the fire resistance time should be within 10 min. However, it should be noted for the designers that when the story drift ratios of CFS structures exceed 2.5% under strong earthquake actions, the fire resistance performance of cavity-insulated load-bearing CFS walls lined with double-layer GP sheathings on either side drop sharply and the residual FRT is only 8min. This implies that once the fire accident occurs after the strong earthquake action, the collapse risk of CFS structure may increase obviously, due to the rapid failure of earthquake-damaged cavity-insulated gypsum-sheathed CFS walls. 4.2. Influence of earthquake damage on the failure modes of the wall studs In the present investigation, structural failure occurs in all four specimens subject to fire experiments, and the following two types of failure modes are generally exhibited by the wall studs. The first failure mode involving specimens W2, W3 and W4 is described as the local buckling of the hot flange approximately 200e700 mm from the top, which further results in a substantial flexural deflect of the stud towards the ambient side (Fig. 19c and d). The mechanism of this failure mode is as follows. The structural failure of the wall stud is generally affected by the following three factors under fire conditions: ① There is a noticeable temperature gradient on the stud section due to the existence of the rock wool cavity insulation, which results in the neutral axis on the stud section (parallel to the flange direction) drifting towards the cold flange. As a

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 22. Postfire visual observation of specimen W5.

result, the additional bending moment generated under the axial force increases the compressive stress of the stud hot flange; moreover, since the temperature on the hot flange of stud is substantially higher than that on the cold flange, the hot flange experiences local buckling first. ② The stud displays out-of-plane deformation towards the inside of the furnace due to the temperature gradient on the stud section. The associated second-order effect decreases the compressive stress of the hot flange; because the secondorder effect near the top of the stud is less than that at the mid-height of the stud, the compressive stress of stud hot flange near the top of the specimen is higher than that of stud hot flange at the mid-height of specimen. ③ The temperature near the top of the specimen is slightly higher than that at other heights of the specimen due to the upward movement of hot air flow (Fig. 24). In other words,

the faster temperature increase near the top of the stud results in faster degeneration of the material properties and greater drift of the neutral axis than that of the mid-height of stud. Meanwhile, the failure location might not be the top end of studs due to the constraints provided by the tracks (Fig. 8a). All these factors lead to the failure point being located in the area of hot flange near the top of the stud for each specimen (Fig. 19d). Then, the first failure mode is generated. The second failure mode occurs for specimen W5 and is described as the local crushing on the stud section approximately 200e400 mm from the stud bottom (Fig. 22c). This failure mode is generated for the following reasons. After experiencing the cyclic loading test, specimen W5 already has local buckling near its bottom. Therefore, in the subsequent fire experiment, the stud will

Fig. 23. Time-temperature profiles of specimens W2, W3, W4 and W5.

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 24. Time-temperature profiles of the No. 4 middle stud in specimens W2~W5.

Table 3 Summary of the failure phenomena in the fire experiments. No.

FRT

Failure criterion

Failure mode

Failure location

Sheathing separation

W2 W3 W4 W5

48 min 41 min 40 min 8 min

Structural failure

Lhot-F Lhot-F Lhot-F LC

300e400 mm 300e600 mm 200e700 mm 2600e2800 mm

No separation

Fig. 25. Time-dependent axial load in the middle stud of each specimen.

Fig. 26. Time-dependent out-of-plane deflection of the No. 4 middle stud in each specimen.

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 29. Comparison of the time-temperature curve of HF for specimens W2, W3 and W4. Fig. 27. Relationship between the drift ratio and FRT of the present CFS walls.

inevitably experience further structural failure at the point where local buckling previously occurred. Hence, for the cavity-insulated load-bearing CFS wall lined on both sides with double layers of GP boards, after being subjected to cyclic loading at a drift ratio within 2%, its failure mode and failure position during postearthquake fire do not change. However, when the drift ratio is greater than 2.5%, local buckling will occur in the area near the bottom of the wall stud; thus, in this case, the failure mode and failure position of wall studs change as a result of the postearthquake fire. 4.3. Influence of the loading conditions on the fire resistance of the CFS wall The present wall specimen W2 has a configuration similar to that of Kodur and Sultans' specimen F38 [15]; however, the loading conditions of the fire experiment are substantially different. In Kodur and Sultans’ fire experiment, the traditional boundary conditions are adopted, i.e., the vertical load is transferred to the wall specimen through the distributive beam, and the vertical load of the wall specimen remains unchanged during the fire experiment. In the present investigation, the vertical load is independently applied to each stud of the wall specimen, and the axial thermal expansion load of the stud caused by the restraint of the rigid floor is considered during the fire experiment. Fig. 31 shows a comparison of the wall steel frame failure modes between Kodur and Sultans' experiment [15] and the present

Fig. 30. State diagram of specimen W5 during fire exposure.

experiment. The failure mode from Kodur and Sultans' specimen is the same as that of the present specimen: local buckling of the hot stud flange, which results in out-of-plane deformation of the wall stud towards the ambient side. However, the failure positions of the previous two specimens are different, and a reasonable explanation for this difference cannot be provided in the present investigation.

Fig. 28. Opening of the fire-side board joints of W2 during the initial stage of fire exposure.

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the code [22], when the axial thermal expansion constraints are considered. Therefore, the investigation of new configurations or new fire-resistant sheathing boards become necessary to improve the fire resistance of such CFS walls. 4.4. Mechanism of cavity insulation on the fire resistance of loadbearing CFS walls Specimen W2 has a similar configuration and load level as our previous specimen without cavity insulation (recorded as Pre2 [11]), wherein the major difference is that specimen Pre2 does not possess cavity insulation. Fig. 32 shows the failure modes of these two specimens. Fig. 33 shows the time-temperature curves and time-dependent out-of-plane deflection curves at the mid-height of these two specimens. The following observations can be noted from these two figures:

Fig. 31. Influence of the conventional and new present loading conditions on the FRTs of CFS walls.

Fig. 32. Influence of the cavity insulation on the failure mode of load-bearing CFS walls.

In addition, the fire resistance time is 59 min for Kodur and Sultans’ specimens, whereas the fire resistance time is only 48 min for specimen W2. The reasons for the shorter fire resistance time of specimen W2 are as follows: ① Each stud of specimen W2 is independently loaded. Therefore, when a stud is damaged, the vertical load cannot be redistributed. ② The axial thermal expansion load of each stud for specimen W2 is very substantial when subjected to fire (in Fig. 25, the maximum axial thermal expansion load of specimen W2 stud reaches 10.6 kN), which will substantially increase the actual axial load level of the stud. Hence, in mid-rise CFS structures, the axial thermal expansion constraints provided by the steel deck floor slabs (one type of rigid floor slab) will reduce the FRT of load-bearing CFS walls and accelerate the structural failure of CFS wall frame under fire conditions. If the effect of axial thermal expansion constraints is not considered, the corresponding fire design of load-bearing CFS walls may become non-conservative. Moreover, it needs special attention in engineering practice that the commonly used cavity-insulated load-bearing CFS walls lined with double-layer GP sheathings on either side do not meet the fire resistance rating of 60 min listed in

(1) The failure modes of the two wall specimens are substantially different, and the failure of the wall stud in specimen Pre2 without cavity insulation is flexural buckling of the stud towards the fire side, which is associated with local buckling on the cold flange at the mid-height (Fig. 32a). This failure mode is also present in the fire experiments of load-bearing CFS walls without cavity insulation conducted by other researchers [13]. The main reason for this failure mode is that the temperature gradient on the stud section of the wall without cavity insulation is substantially smaller than that of the cavity-insulated wall; hence, the drift of the neutral axis, which is parallel to the flange direction, in the previous noncavity insulated stud section is smaller than that in the current cavity-insulated stud section, whereas the second-order effect caused by thermal bowing plays an important role in exacerbating the out-of-plane deformation (Fig. 33b) towards the fire-side and the buckling of the stud. The failure mode of the cavity-insulated wall is local buckling of the hot flange approximately 200e700 mm from the top, which further results in substantial flexural deflection of the stud towards the ambient side (Fig. 32b). The mechanism of this failure mode has been introduced in Section 4.2 and will not be repeated here. (2) The FRTs of the two specimens are substantially different: 48 min and 71 min for specimens W2 and Pre2, respectively. This difference means that the FRT of the load-bearing wall without cavity insulation is substantially higher than that of the load-bearing wall with cavity insulation. This phenomenon is consistent with the results of the fire experiments on load-bearing walls CFS conducted by Kodur and Sultan [15] and Kesawan and Mahendran [18]. Previous studies have explained that the existence of cavity insulation hinders the heat transfer to the ambient side, accelerates the temperature increase in the GP board on the fire side and promotes its cracking and at least partial fall-off, resulting in a rapid temperature increase and premature local buckling on the hot flange of the steel stud, which further leads to the loss of bearing capacity in the CFS wall. However, this explanation is not completely accurate because none of the present loadbearing wall specimens (W2~W5) experienced partial falloff of the GP board on the fire side, and the rapid temperature increase of the fire-side GP board will inevitably lead to a fast temperature increase on the directly contacted hot flange of the stud. Therefore, a more concise description is that the existence of the cavity insulation hinders the heat transfer from the cavity to the ambient side, which results in a rapid temperature increase in the fire-side sheathing and the hot stud flange; this phenomenon leads to the temperature in the CFS wall with cavity insulation being

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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Fig. 33. Influence of the cavity insulation on the time-dependent curves of load-bearing CFS walls.

substantially higher than that in the CFS wall without cavity insulation (Fig. 33a). As the neutral axis (which is parallel to the flange direction) drifts towards the cold flange and material properties of the hot flange rapidly degenerate, local buckling of the hot flange on the steel stud occurs first in the cavity-insulated load-bearing CFS wall, which further results in the substantial flexural deflection towards the ambient side. As a result, for the design practice of CFS structures, it is a common sense that the existence of cavity insulation will improve the sound and heat insulation performance of load-bearing CFS walls. Unfortunately, the existence of cavity insulation becomes a disadvantage configuration to the fire resistance of gypsumsheathed load-bearing CFS walls. It will not only reduce the FRT, but also may change the failure mode of CFS wall studs under fire conditions, due to the rapid temperature increase of stud hot flange and significant temperature gradient between stud hot and cold

flange. Hence, it is worth developing a new configuration of CFS load-bearing wall, which can have good fire-resistance, sound and heat insulation performance, and is convenient for construction. 5. Conclusions This paper reports a detailed postearthquake fire experimental investigation of full-scale cavity-insulated load-bearing CFS walls lined on both sides with double-layer GP boards. The conclusions of this study are as follows. (1) For the design practice of CFS structures, when the story drift ratios of CFS structures are less than 2.0% under the design earthquake actions, the residual FRT of cavity-insulated loadbearing CFS walls lined with double-layer GP sheathings on either side is only 8 min less than that of wall assemblies without earthquake damage. Meanwhile, the failure mode of CFS wall frame is local buckling of the hot flange near the top

Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845

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of the stud, leading to the substantial flexural deflection of the stud towards the ambient side. (2) It needs special attention for the designers that when the story drift ratios of CFS structures exceed 2.5% under strong earthquake actions, the fire resistance performance of cavityinsulated load-bearing CFS walls lined with double-layer GP sheathings on either side drop sharply and the residual FRT is only 8min. Meanwhile, the failure mode of CFS wall frame is changed to local crushing near the bottom of the stud under fire conditions. Hence, once the fire accident occurs after the strong earthquake action, the collapse risk of gypsumsheathed CFS structure may increase obviously. (3) In mid-rise CFS structures, the effect of axial thermal expansion constraints provided by the steel deck floor slabs (one type of rigid floor slab) to the CFS wall studs is significant under fire conditions. It will accelerate the structural failure of load-bearing CFS walls and should be considered in the fire design of load-bearing CFS walls. Moreover, it should be noticed that the commonly used cavity-insulated loadbearing CFS walls lined with double-layer GP sheathings on either side do not meet the fire resistance rating of 60 min, when the axial thermal expansion constraints are considered. (4) The existence of cavity insulation is a disadvantage configuration to the fire resistance of gypsum-sheathed loadbearing CFS walls. It will not only reduce the FRT, but also may change the failure mode of CFS wall studs under fire conditions, due to the rapid temperature increase of stud hot flange and significant temperature gradient between stud hot and cold flange.

Conflicts of interest The authors declared that they have no conflicts of interest to the present investigation. Acknowledgments This research was supported by the National Natural Science Foundation of China (Grant No. 51538002, No.51508088). Appendix A. Supplementary data Supplementary data related to this article can be found at https://doi.org/10.1016/j.jcsr.2019.105845. References [1] K. Roy, J.B.P. Lim, H.H. Lau, P.M. Yong, G.C. Clifton, A. Wrzesien, C.C. Mei, Collapse behaviour of a fire engineering designed single-storey cold- formed

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Please cite this article as: W. Chen et al., Postearthquake fire experiments of cavity-insulated cold-formed steel load-bearing walls lined with plasterboards, Journal of Constructional Steel Research, https://doi.org/10.1016/j.jcsr.2019.105845