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Load-slip behaviour of steel-cross laminated timber (CLT) composite connections
Journal of Constructional Steel Research 122 (2016) 110–121
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Journal of Constructional Steel Research
Load-slip behaviour of steel-cross laminated timber (CLT) composite connections A. Hassanieh, H.R. Valipour ⁎, M.A. Bradford Centre for Infrastructure Engineering and Safety School of Civil and Environmental Engineering, UNSW Australia, Sydney, NSW 2052, Australia
a r t i c l e
i n f o
Article history: Received 1 December 2015 Received in revised form 3 March 2016 Accepted 9 March 2016 Available online xxxx Keywords: Cross laminated timber (CLT) Load-slip Shear-connectors Steel-timber composite (STC)
a b s t r a c t Connecting cross laminated timber (CLT) panels to steel girders using mechanical connectors (e.g. screws and bolts) is an attractive and novel method for developing a fully prefabricated and sustainable hybrid steeltimber composite (STC) floor that can also facilitate future dismantling and recycling of its structural components. Since the structural performance of composite floors (including STC systems) is significantly influenced by strength and stiffness of slab-to-girder composite joints, this study characterises the load-slip behaviour and failure modes of steel-CLT timber composite joints by conducting push-out tests on three different types of STC connections with high-strength bolts, coach screws or a combination of glue and coach screws. Furthermore, the effect of reinforcing the CLT slabs by using steel nail plates on the strength and stiffness of STC joints is evaluated. Empirical models for capturing load-slip behaviour of steel-CLT composite joints with dowel (i.e. screw and bolt) connectors are derived from non-linear regression of experimental data and simple formulae for the strength and stiffness of steel-CLT composite joints with dowel connectors are proposed. Moreover, 1D beam on inelastic foundation finite element (FE) models of the dowel STC connectors are developed and validated against experimental results. © 2016 Elsevier Ltd. All rights reserved.
1. Introduction From a structural engineering perspective, replacing reinforced concrete (RC) slabs with timber flooring can reduce the self-weight of a building significantly, leading to a reduction in energy-intensive concrete foundations as well as to less-demanding rigging and craneage requirements. In addition, lowering the self-weight of the structure significantly reduces the inertial forces induced by seismic actions, particularly on medium to high-rise buildings, and it also facilitates the construction of buildings on soft and/or problematic soils. In steeltimber composite (STC) floors, bolts and self-tapping screws can be utilised to transfer the interface shear between the steel joist and the prefabricated timber slab. STC composite joints with mechanical connectors (i.e. screws and bolts) can also facilitate the dismantling, recycling and reuse of structural components. Cross laminated timber (CLT) is a widely-used innovative building material fabricated by bonding together timber boards to produce a solid panel, with each layer of the panel alternating between longitudinal and transverse lamellae. Owing to the alternation of the grain directions, CLT panels have improved dimensional stability and a much lower variability of their mechanical properties when compared with solid timber. Furthermore, CLT panels have relatively high strength and
⁎ Corresponding author. E-mail address: [email protected] (H.R. Valipour).
stiffness properties in both the longitudinal and transverse directions that make possible the two-way spanning action needed to resist transverse loading. Accordingly, CLT panels are an attractive candidate for developing hybrid STC floors/beams. The structural performance and behaviour of steel-CLT composite floors is significantly influenced by mechanical properties of the CLT panels and the strength and stiffness of composite joints. Several studies have been conducted hitherto to characterise the mechanical properties (tensile, compressive, bending and embedding strengths) of CLT. For example, embedding tests have been performed on CLT panels according to the specifications of EN383-2007 [1], and the results have been used to determine the load carrying capacity of CLT joints [2]. Furthermore, the behaviour and mechanical properties of CLT under compressive loading perpendicular to the grain direction and the effect of boundary line loading have been investigated by Serrano and Enquist [3], and it was shown that the brittle compressive failure mode of CLT panels depends on the loading arrangement. Apart from standardised tension, compression and embedding tests, plate-bending tests on three- and five-layered CLT panels have been carried out to determine the global and local failure mechanisms and the evolution of different crack modes in the CLT lamellae [4]. Based on the plate bending tests, the ductile structural behaviour of CLT plates was also described [4]. In addition to plate bending tests, an advanced laminated plate theory was employed by Stürzenbecher et al. [5] to accurately capture the mechanical behaviour of CLT panels under distributed and concentrated loading.
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Over the past decade, several experimental and theoretical studies have focused on the behaviour of CLT panel-to-panel connections. The load carrying capacity and ductility of double lap CLT joints with dowel connectors have been studied by Blaß et al. [6]. Large embedding deformations were observed in the CLT lamellae, and using punched metal plate fasteners or glued-on boards was found to be effective in improving the embedding strength of the CLT panels as well as the load carrying capacity of double lapped CLT joints [6]. Moreover, monotonic and cyclic shear tests on lapped and orthogonally connected CLT/X-Lam panels (representing walls or floors) with screws and steel angle brackets were performed by Gavric et al. [7,8] to evaluate the strength, stiffness, energy dissipation and ductility of CLT/X-Lam wall-to-wall, wall-to-floor/foundation and floor-to-floor connections. Similarly, the cyclic/seismic behaviour of CLT shear wall-to-floor connections with steel brackets and mechanical fasteners in both parallel and perpendicular to the grain directions has been investigated experimentally and theoretically by Shen et al. [9] and Schneider et al. [10]. The results of cyclic tests conducted by different researchers have demonstrated the reasonable seismic performance of CLT shear walls with mechanical connectors such as nails and screws [10,11]. In addition to behaviour of CLT panel-to-panel connections, several studies have been performed experimentally and numerically on the seismic performance of CLT structures, and the influence of connections' behaviour on the global response of the buildings with CLT shear walls and floors [8,12–18]. However, experimental data on behaviour of steel-timber and concrete-timber composite floors with CLT slabs and dowel connectors is very limited. Very recently, Loss et al. [19] performed symmetric push-out tests to characterise the behaviour of different connections in steel-CLT composite floors [19]. Furthermore, the load-slip behaviour of timber-timber and concrete-timber composite connections with CLT slabs have been investigated by conducting push-out tests on asymmetric specimens [20,21]. This paper investigates the short-term load-slip, peak load carrying capacity, stiffness and failure mode of steel-CLT composite joints by conducting symmetric push-out tests. In the specimens tested, the CLT panels are connected to the flanges of the steel profile using bolts, coach screws and/or glue. Non-linear regression of the experimental results is undertaken to derive empirical load-slip formulae for steel-CLT composite joints with coach screws and high-strength bolts. The empirical load-slip models developed can be incorporated in finite element models and used for the nonlinear analysis of steel-CLT composite beams/floors with dowel connectors. In addition, 1D beam on inelastic foundation finite element (FE) models of the screw connectors are developed and validated against experimental push-out tests.
2. Experimental program 2.1. Specimen details The push-out tests were conducted on nine different groups of symmetric steel-CLT specimens that comprise of two CLT panels connected to the flanges of a steel profile (Fig. 1). The specimens were categorised with respect to the type of connectors (i.e. bolts and screws with and without glue) and application of nail plate fasteners for reinforcing the CLT panels. Furthermore, an end connection was designed to be used at the two ends of the simply supported composite beams. The end connection comprised of a 50 × 50 × 5 mm steel box connected to the flange of the steel girder by two high-strength friction grip Grade 8.8 bolts that were prestressed to 60% of their yield load (Fig. 1). The geometry, details of the specimens, test set up, location of loads and supports and type of dowel connectors are shown in Fig. 1 and in Table 1. The primary variables within the different push-out specimens were the
Fig. 1. Configuration and cross-sections of nine different steel-CLT composite joints for push-out tests.
diameter and type of the mechanical connectors (i.e. bolts or screws) and the use of glue or punched steel plate fasteners in conjunction with screws.
Table 1 Size and configuration of connectors and loading direction in the push-out specimens. Group
Mechanical connectors
With nail plate Without nail pate Glued End connection
Screw 12, 16, 20 & bolt 12, 16 Screw 12, 16 Glue + screw 16 50 × 50 × 5 box connection + bolt 12 + screw 12
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2.2. Material properties 2.2.1. Cross laminated timber (CLT) The CLT panels were 120 mm thick and they were made from five layers (Fig. 2a) of Spruce lamellae of strength class C24 according to BS EN338 [22]. The orientation and thickness of each lamella are shown in Fig. 2a. The CLT panels were 400 mm wide and 600 mm long and in the push-out tests the CLT panels were loaded in the direction parallel to the grain of the first layer (Fig. 1). The CLT panels had an average moisture content (MC) of 12% (±2%) and density of 490 kg/m3 at the time of fabrication and testing of the STC specimens. The nominal
mechanical properties of the CLT panels (at 12% MC) are given in Table 2. The mechanical properties of timber/engineered wood in different standards such as ASTM D2555 [23] and ASTM D3737 [24] are typically given as ratios of mechanical properties at a specific MC, i.e. 12%. It is well-established that mechanical properties and behaviour of timber and engineered wood products can be influenced by the moisture content and its variations [25,26]. Generally, the strength and elastic modulus of timber decline as the moisture content of timber increases. This change in timber modulus of elasticity and/or strength can affect the load-slip and peak load capacity of the STC system. 2.2.2. Steel profile The push-out specimens were fabricated by connecting CLT panels to Australian 310UB32 hot rolled I-sections that comply with the specifications of AS/NZS 3679.1 [27]. The dimensions, yield strength and ultimate strength of the 310UB32 cross-section are given in Table 3. 2.2.3. Dowel connectors Hexagonal high-strength Grade 8.8 bolts having diameters of 12 mm and 16 mm, and hexagonal coach screws having diameters of 12 mm, 16 mm and 20 mm were used in the fabrication of the specimens (Fig. 2b). The screws were made of Grade 4.6 steel with a yield strength of 240 MPa and an ultimate tensile strength of 400 MPa, and they comply with the minimum requirements of AS/NZS 1393 [28]. The highstrength bolts had a proof yield strength of 660 MPa and an ultimate tensile strength of 830 MPa, as per the specifications of AS1110.1 [29] and AS1112.1 [30]. The coach screws were 100 mm long and the bolts were 110 mm long. 2.2.4. Nail plate The specimens with 12 and 16 mm screw connectors were tested twice, i.e. with and without nail plates. The nail plates were used to reinforce the CLT panels around the connector hole and to improve the stiffness and strength of the STC joints. The nail plates were 100 × 50 × 1 mm (Fig. 2c) and made of Grade G300 steel with a characteristic yield strength of 300 MPa and a tensile strength of 340 MPa. In the composite joints with 12 and 16 mm screws, one and two nail plates per screw were used respectively to reinforce the CLT panels. The nail plates were installed on the CLT panels simply by hammering. After installation of the nail plates, the holes required for installation of dowel fasteners (i.e. bolts or screws) were drilled in the nail plates and then the CLT panels were mounted on steel beam flange and the coach screws/bolts were installed. Only one nail plate was used in conjunction with each 12 mm shear connector, however, two nail plates installed in a cruciform configuration were used in conjunction with 16 mm dowel connectors. 2.2.5. Glue In some of the specimens, a non-sag gel type epoxy was used in conjunction with 16 mm coach screws to provide full composite action between the CLT panels and steel profile. The mechanical properties and curing requirements of the epoxy are provided in Table 4. 2.3. Test setup and instrumentation of the specimens
Fig. 2. (a) 5-Layer cross laminated timber (CLT) panel (b) bolt and coach screw connectors and (c) nail plate.
A symmetric CLT-steel-CLT arrangement (Fig. 1) for the fabrication of specimens and test set up was adopted in this study to prevent unsymmetrical load distribution on the sample and to minimise unwanted friction between different parts of the specimen [31]. The symmetric configuration of the specimens can provide a close to uniform distribution of load with predictable friction effects on the load-slip response, and it can also facilitate fabrication of the push-out test specimens. The relative displacement (or slip) between the CLT panels and steel profile was measured using four linear variable differential transformers (LVDTs) having 100 mm maximum stroke. Furthermore, an LVDT was used to measure the vertical movement of testing machine actuator as
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Table 2 Mechanical properties of CLT (in MPa). Bending (fb)
Tension parallel to grain (ft)
Shear in beams (fs)
Compression parallel to grain (fc)
Compression perpendicular to grain (fp)
Elastic modulus (E)
Modulus of rigidity (G)
24
16.5
4.6
24
2.7
12,000
690
Table 3 Mechanical and geometrical properties of steel profile. Section depth d (mm)
Flange width bf (mm)
Flange thickness tf (mm)
Web thickness tw (mm)
Yield strength fy (MPa)
Ultimate strength fu (MPa)
Elastic modulus Es (MPa)
298 mm
149 mm
8 mm
5.5 mm
320 MPa
450 MPa
205 GPa
well as the specimens. A steel-CLT composite specimen placed in the testing rig and location of LVDTs mounted on the specimen are shown in Fig. 3. One of the main objectives of this study was to develop a sustainable and demountable steel-CLT composite floor with respect to practical construction considerations. Accordingly, the CLT panels and flanges of the steel profiles were predrilled and the specimens were then assembled following a procedure similar to that representative of current construction practice. The dimeter of the holes predrilled in the flanges of the steel profile was 0.2 mm larger than the dimeter of the screws/ bolts. The diameter of the holes predrilled in the CLT panels was 2 mm smaller than the diameter of the screws, and 1 mm to 2 mm larger than the diameter of the bolts. In the bolted connections, a 50 mm × 50 mm washer was used to prevent crushing of the CLT in the direction perpendicular to the plane of the timber lamellae. A calibrated torque wrench was used to tighten the high-strength bolts to induce a post-tensioning force of 10 kN equivalent to 0.14fy in the 12 mm bolts and 0.08fy in the 16 mm bolts (fy being the yield stress of the bolt). The post-tensioning force was calculated based on a design crushing strength of approximately 4 MPa for CLT and area of 2500 mm2 for the square washers. In the glued steel-CLT composite joints, the epoxy adhesive was applied over a 450 mm by 152 mm area on the surface of the CLT panels and after mounting the CLT panels, the specimen was left for one hour and the 16 mm screws were then installed. The thickness of the adhesive layer after drying was around 2 mm.
3. Discussion of test results 3.1. Modes of failure Four distinctive modes of failure, identified as Modes I, II, III and IV, were observed in the push-out tests on timber-timber, timberconcrete and timber-steel joints with dowel connectors [33–36]. These modes of failure have been also recognised in timber structures standard BS EN 26891:1991 [32]. The occurrence of each mode of failure depends on the slenderness of the connectors represented by L/D (with L being the length and D the diameter of the connector) and the strength ratio fs/bs (with fs being material strength of the connector and bs the bearing strength of the primary element [35]). Failure Mode I is associated with crushing of timber in the vicinity of the dowel connectors without significant deformation in the dowels (Fig. 5). In failure Mode II, a plastic hinge forms within the dowel and crushing of timber occurs close to the surface of the timber element, whilst in Mode III, crushing of the timber is associated with the formation of two plastic hinges, one at the middle and the other one close to the end of the connector (Fig. 5). Failure mode IV is associated with fracture of the connectors and significant crushing of the timber. It is noteworthy that failure Modes I and IV are brittle, however, failure Modes II and III can be ductile or semi-ductile depending on L/D and fs/bs. The dominant failure mode of the steel-CLT composite joints with 12 mm screws (without a reinforcing nail plate) that were tested was Mode III, while in the composite joints with 16 mm and 20 mm screws (without a reinforcing nail plate), the failure mode was Mode II (Fig. 6a– c). In steel-CLT composite joints with bolted connectors (including the
2.4. Loading procedure The specimens were loaded following a procedure specified in the Eurocode for timber structures [32]. In the first stage, the load was increased from 0 to 0.4Fu (with Fu being the estimated ultimate load) over 120 s and the load was maintained at 0.4Fu for 30 s (Fig. 4). The second stage involved unloading of the specimen from 0.4Fu to 0.1Fu and then maintaining the load at 0.1Fu for 30 s. In the final stage, testing was conducted under a load control regime up to 0.7Fu, and then reverted to displacement-control (with an average displacement rate of 2 mm/s) and continued until failure of the specimen (Fig. 4). The loading procedure in each test was redefined with respect to the ultimate load Fu observed in the previous test.
Table 4 Mechanical properties and curing requirements of the epoxy adhesive. Tensile strength
Compressive strength
Shear strength
Coefficient of linear expansion
Minimum cure time at 25 °C
30 MPa
70 MPa
15 MPa
60 × 10−6 mm °C
24 h
Fig. 3. Push-out test setup and instrumentation.
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Fig. 4. BS EN 26891 loading procedure adopted for the push-out tests [32].
joints with end 50 × 50 × 5 mm steel box connections), Mode IV was the dominant failure mode (Fig. 6d). Using nail plates to reinforce the CLT panels did not alter the failure mode of the joins with screw connectors, but the nail plates reduced the ductility of joints slightly. The reduction in the ductility of joints with nail plates was evident from the sudden drop of the load after the peak load was attained. The failure mode of joints with reinforcing nail plates was associated with rupture of nail plate, localised crushing of the timber and the formation of one or two plastic hinges in the screws, depending on the dimeter of the screws (Fig. 7a). In general, the joints with screw connectors exhibited a relatively ductile behaviour, whereas the joints with bolted connectors or combination of glue and screws produced a brittle mode of failure that was associated with fracture of the bolts or glue (see Fig. 7b) and separation of the CLT panels from steel profile. It is noteworthy that the steel profiles and the bolt or screw holes in the flange of steel profile did not exhibit any sign of deformation during or after the test. 3.2. Load-slip response of connections It is well known that the variability of timer mechanical properties and the method of fabrication can significantly affect the structural behaviour of timber connections. Accordingly, three identical specimens
Fig. 5. Characteristics of failure modes I, II and III observed in the steel-CLT composite joints.
Fig. 6. Failure mode of steel-CLT composite joints with (a) screw S12, (b) screw S16 and (c) bolt B12 and (d) bolt B16 connectors.
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Fig. 7. (a) Crushing of timber and nail plate rupture in the steel-CLT composite joints with screw connectors and reinforcing nail plate and (b) fracture in the glued joints.
Fig. 8. Load-slip response of steel-CLT composite joint with (a) 2 × S12 (12 mm screw) (b) 2 × S16 (16 mm screw) and (c) 2 × S20 (20 mm screw).
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were fabricated for each type of steel-CLT composite joint (Table 1 and Fig. 1), and the specimens were tested under identical loading conditions and test set up, to ensure the precision and reproducibility of the experimental data (i.e. load-slip response, stiffness under service load condition, peak load capacity and failure mode). The load-slip responses of the STC joints with screws (with and without a reinforcing nail plate) and bolted connectors obtained from the three identical specimens are shown in Figs. 8 to 10 respectively. Furthermore, the mean load-slip curves are also shown in Figs. 8 to 10, which demonstrate clearly the small variation in the peak load capacity and of the load-slip response of the steel-CLT composite joints. In the bolted steel-CLT composite joints, the friction between the steel profile and the CLT panel mobilised by the post-tensioning of the high strength bolts has provided full composite action at the early stages of the load-slip evolution, where no slip occurs up to a shear force of 11 kN (Fig. 9). The load corresponding to first slip (see Fig. 9) in the bolted STC joints depends on the magnitude of post-tensioning force induced in the high-strength bolts. Accordingly, increasing the initial post-tensioning force in the bolted connectors can increase the initial stiffness (initial slope of the load-slip behaviour) of the bolted connectors, however, the magnitude of post-tensioning force has minor influence on the peak load carrying capacity of the bolted STC joints. It is noteworthy that the magnitude of post-tensioning force in the bolts must be specified/limited with respect to the compressive strength of CLT panels to ensure that crushing/punching (e.g. Fig. 6d) of CLT in the direction perpendicular to the plane of lamella would not happen.
Fig. 9. Load-slip response of steel-CLT composite joint with (a) 2 × B12 (12 mm bolt) and (b) 2 × B16 (16 mm bolt).
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Fig. 10. Load-slip response of steel-CLT composite joint with reinforcing nail plates and (a) 2 × S12 (12 mm screw) and (b) 2 × S16 (16 mm screw) connectors.
Apart from steel-CLT joints with dowel (i.e. screw and bolt) connectors, a composite joint with a combination of glue and screws and also a composite joint with friction grip end 50 × 50 × 5 mm steel box connections were fabricated and tested. The experimental load-slip and shear stress-slip results for the glued joints and the experimental load-slip response of the composite joints with friction grip end 50 × 50 × 5 mm steel box connections are shown in Figs. 11 and 12, respectively. Moreover, the mean and coefficient of variation (CoV) of the peak load carrying capacity and the mean of the serviceability stiffness (represented by the slope of the load-slip curve between 10% and 60% of the peak load capacity, i.e. the ks,0.6 slip moduli) for the steel-CLT composite joints are given in Table 5. The maximum CoV of the peak load carrying capacity of the steel-CLT composite joints with screw and bolt connectors was limited to CoVpeak = 8.8% and 3.9%, respectively. With regard to the load-slip response and peak load capacity of the tested composite joints (Figs. 8 to 12 and Table 5), the following conclusions can be established: • The glued steel-CLT composite joints exhibited significantly higher strengths and stiffnesses compared to the joints with only mechanical connectors (i.e. dowelled or friction grip end box connections). Furthermore, the friction grip end box connections can have higher peak load carrying capacity and stiffness than conventional dowelled connectors. • The load carrying capacity of steel-CLT joints with bolted connectors is higher than for joints with screws, because of the higher strength of the steel used for manufacturing the Grade 8.8 bolts compared to the mild steel used for manufacturing the screws. • The experimental load-slip responses of the joints with screw connectors (Figs. 8 and 10) exhibited close to elastic-perfectly plastic behaviour, whilst the steel-CLT joints with bolted connectors showed some hardening characteristics (see Fig. 9).
Fig. 11. (a) Load-slip and (b) stress-slip response of glued steel-CLT composite joints.
• Reinforcing the CLT panels by nail plates can enhance the load carrying capacity and particularly the serviceability stiffnesses of steelCLT composite joints with screws. A minimum strength enhancement of 28% was achieved in composite joints with S16 screws where two nail plates per S16 screw hole were used to reinforce the CLT panel. Furthermore, a minimum stiffness increase of 70% for S12 and 226% for S16 connectors was observed. However, reinforcing the CLT panels by nail plates can reduce the ductility of steel-CLT composite joints. The less ductile behaviour of joints with nail plates is evident from the post-peak softening of load-slip response (see Fig. 10). 4. Analytical load-slip model for steel-CLT composite joints with dowel connectors In this section, empirical load-slip models for the tested steel-CLT composite joints with dowel connectors are developed by non-linear
Fig. 12. Load-slip response of the steel-CLT composite joints with end 50 × 50 × 5 mm steel box connection.
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Table 5 Mean and CoV (in %) of the peak load capacity and mean of the slip moduli ks,0.6a for the STC joints. Screwb
Nail plate Slip moduli ks,0.6 (kN/mm) Peak load capacity (kN)
Without With Without With
Boltb
S12
S16
S20
B12
B16
9.52 16.23 35.78 [3.7] 38.88 [6.3]
8.21 26.8 49.87 [8.8] 63.98 [3.8]
6.11 – 67.52 [2.4]
6.26 – 79.88 [2.2]
9.66 – 123.4 [3.9]
Box 50 × 5 × 5 + 12 mm bolt 34.94 – 131.13 –
a The slip moduli ks, 0.6 is used as a representative of the connection stiffness within the service range of loading. The ks, 0.6 slip moduli is the slope of the load-slip curve between 10% and 60% of the peak load capacity. b The values in [] are CoV (in %) obtained from test results on three identical specimens.
regression analysis of the push-out test results. The proposed load-slip function is inspired by Ramberg-Osgood [37] model that has been widely used in analysis of structural steel components [38]. Essentially, a post-peak softening branch has been incorporated into the RambergOsgood model to represent both hardening and softening behaviour of steel-timber composite connections with just one mathematical function. The proposed model can be easily incorporated into componentbased finite element models and used for the non-linear analysis of hybrid steel-CLT composite beams/floors with coach screw and bolt connectors. For the sake of ease of computer implementation, a seven-parameter model built on three asymptotic lines (see Fig. 13) is proposed, being given by k0 −kp s kp þ ks s þ f ¼n 1 h h i n2 on1 −ks s; i n1 on1 n 2 1 þ k0 −kp fs 1 þ kp þ ks f −s f 0
1
ð1Þ
0
where f is the shear force, s the slip, k0 the initial stiffness, kp the prepeak stiffness, ks the post-peak stiffness, f0 the first reference shear force corresponding to the pre-peak branch and f1 the second reference shear force corresponding to the post-peak branch, and where n1 and n2 are two parameters that control the curvature of the first and second curves, respectively (Fig. 13). It is noteworthy that the stiffness values
k0, kp and ks can be positive, negative or zero and accordingly the proposed empirical model can capture the pre-peak as well as the postpeak load-slip response of the steel-CLT composite joints with dowel connectors. The non-linear regression based on the least squares method with trust-region algorithm and bisquare robustness available in MATLAB software was employed to fit the proposed function to the experimental data and determine the values of seven different input parameters, i.e. k0, kp, ks, f0, f1, n1 and n2. In the first step, the mean experimental loadslip for each group of steel-CLT joints was obtained. In the second step, the initial values and upper and lower bounds for the seven unknowns are defined with respect to the mean of experimental loadslip and the values of each parameter within a 95% confidence bound is calculated using the adopted non-linear regression procedure. The seven input parameters required for establishing the analytical loadslip model of the steel-CLT composite joints with mechanical connectors (i.e. screw, bolt and end box friction grip connection) are given in Table 6. In addition, the R-square values for the calibrated load-slip curves are provided in Table 6. It is seen that the R-square values for all calibrated models are greater than 0.9798, demonstrating goodness-of-fit and close correlation between the analytical model and experimental data. The excellent correlation between the proposed analytical models and the experimental results is also evident from the load-slip curves shown in Fig. 14. In addition, a linear regression was carried out on the experimental slip moduli ks,0.6 and peak load carrying capacity of the joints with coach screw connectors versus the diameter of screws and the results are shown in Fig. 15. Accordingly, a formula is proposed for calculating the ultimate load carrying capacity Pu as 2
P u ¼ 0:123d þ 18:28
ð2Þ
and the serviceability stiffness (represented by slip moduli ks,0.6) of the steel-CLT composite joints with screw connectors as 2
ks;0:6 ¼ 3:94d −11:96
Fig. 13. Proposed analytical model for load-slip behaviour of steel-timber composite joints.
ð3Þ
where, Pu is the ultimate strength of peak load carrying capacity (in kN), d the screw dimeter (in mm) and ks,0.6 the slip modulus (in kN/mm).
Table 6 Input parameters for the analytical load-slip model of the steel-CLT composite joints with mechanical connectors (i.e. screw, bolted and end box friction grip connection). Connector
f0 f1 k0 kp ks n1 n2 R2
Screw without nail plate
Screw with nail plate
Box 50 × 50 × 5 + 12 mm bolt
S12
S16
S20
B12
Bolt B16
S12
S16
B12
26.7 42.4 47.7 0.741 0.337 1.03 8.3 0.9985
49.5 284.5 198.8 1.569 0.778 0.5 0.52 0.9943
52.4 73.8 310.7 1.836 ‐0.069 0.45 9.1 0.9998
6.8 40.9 114.5 11.3 -1.396 8.48 2.75 0.9987
5.5 69.8 122.1 16.27 -1.097 2.24 3.35 0.9998
38.2 85.5 41.5 0.144 1.654 1.19 4.015 0.9997
37.4 65.5 64.5 52.8 0.553 6.294 2.13 0.9798
11.1 301 372 52 9.6 5 1.14 0.998
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Fig. 15. Correlation between coach screw diameter and mean of experimental (a) peak load capacity and (b) slip moduli ks,0.6.
(Fig. 16a). The screws are modelled by a non-linear force-based beamcolumn element with distributed plasticity which is available in the OpenSEES [45] library. The foundation behaviour is modelled by zerolength elements that connect the nodes on the beam (connector) to fixed supports by springs with predefined force-displacement relationships (Fig. 16a). For modelling friction between the screw and timber along the axis of screw, another zero-length element was used. The behaviour of the springs used for capturing the friction effect was obtained
Fig. 14. Correlation between analytical model and mean of experimental load-slip response for steel-CLT composite joints with (a) screw connectors, (b) bolted connections and (c) screw connector and reinforcing nail plate.
5. Modelling the connectors using beam on (in)elastic foundation theory The theory of beams on (in) elastic/inelastic foundation can be used to model and predict the behaviour of shear connectors in concrete or timber [35,39,40]. In this modelling strategy, the shear connector (i.e. bolt, screw and nail) is considered as a beam placed on a foundation made of concrete or timber. Accordingly, several studies have used 1D beam elements in conjunction with linear/non-linear springs (representing the foundation) to capture the behaviour of shear connectors in timber-concrete, timber-timber and steel-timber composite joints [41–44]. The main input parameters required for developing such models are the yield strength of the connectors and the embedding stiffness and strength of timber that can be obtained from embedding test results [43]. In this study, the Beam-on-Foundation model (BFM) is employed to capture the behaviour of steel-CLT composite joints with coach screws
Fig. 16. (a) BFM model (b) deformed connection.
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Fig. 17. Adopted uniaxial stress-strain relationship for steel.
from withdrawal tests [2] and the withdrawal capacity R (in N) of the axially loaded screws was obtained from 0:8 0:9 lef
ð4Þ
R ¼ 31 d
where d is the screw diameter (in mm) and lefis effective penetration length (in mm). The effective penetration length lef is obtained by subtracting the length of the cone shaped part of the screw shank (approximately 5 mm) and the steel profile flange thickness (8 mm) from the total length of the screw. It is noteworthy that the effective penetration length can have some influence on the load-slip behaviour and peak load capacity of the screw connectors that requires further investigation. In the FE model, it is assumed that the steel profile flange can provide a rotationally fixed support for the connectors (Fig. 16), owing to its significantly larger flexural stiffness compared to the flexural stiffness of the connectors (screws). To capture the yielding of steel, the screw (beam) cross-sections were discretised to layers/fibres and a uniaxial constitutive law was assigned to each layer/fibre. A multilinear stress-strain relationship (see Fig. 17) was used for modelling Grade 4.6 steel. Moreover, the behaviour of timber foundation was characterised using the results of embedding test performed by Nakashima et al. [46] on CLT specimens. The embedding strength versus displacement curves in Nakashima et al. [46] were converted to force-displacement curves for foundation springs in both parallel and perpendicular to the grain direction. The wood foundation parameters and embedding stress versus displacement curves adopted for developing the force-displacement relationship of the foundation springs are provided in Table 7 and Fig. 18, respectively. It is noteworthy that the layer-wise characteristics of CLT panels were considered in the beam-on-foundation finite element models by assigning force-displacement relationships in parallel and perpendicular to the grain directions to different parts of the connector as shown in Fig. 16. The load-slip response of the steel-CLT composite joints with screw connectors predicted by the beam on foundation FE models are compared with the experimental results in Fig. 19 and a reasonably good correlation between the FE model predictions and experimental data (particularly the initial stiffness and peak load capacity) is observable. However, the yield load of the composite joints predicted by the FE Table 7 Adopted wood foundation parameters. f0l MPa
f0y MPa
f90l MPa
f90y MPa
d0l mm
d0y mm
d90l mm
d90y mm
α0
α90
21
33
9
12
0.6
1.31
0.78
1.21
0.01
0.05
Fig. 18. Adopted wood foundation stress-displacement behaviour for (a) parallel and (b) perpendicular to the grain direction.
model has only a mediocre accuracy. The discrepancy between the FE and experimental results can be attributed to variability of steel and timber mechanical properties with respect to diameter of screws (size effect) and also the influence of drilling and installation methods on the mechanical properties of timber in the vicinity of the connector holes. 6. Conclusions The study illustrated the general behaviours of various connection types in CLT-to-Steel composite parts as the most critical element in timber-steel hybrid structures. The main points of the study are the following: • The load-slip response of coach screw connections demonstrated ductile behaviour, and typically failure occurs after a large post-peak branch in the load-slip curve. Local timber crushing and developing plastic hinges in coach screws causes failure of whole specimen. The pre-stressed bolted connections show a ductile response and failure is associated with rupture and sudden drop in load slip response. • Glued connections are much stiffer and stronger than screw and bolted connections, but they have a brittle failure that undesirable. Using glued connections in conjunction with the coach screws leads to more ductile behaviour of the connection.
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References
Fig. 19. Comparison of load-slip response predicted by beam on foundation FE models for (a) 2 × S12, (b) 2 × S16 and (c) 2 × S20 connectors.
• Reinforcing the coach screws connections with nail plates can increase the initial stiffness and strength of the connections. • The end connection demonstrated considerably higher stiffness and strength. This connection can be used in beam/floors where positive bending moment is applied on these elements. Although it is considered as the first and last shear connections in composite beams, it may be used in the first and last spots of each panel segment with some considerations. • The analytical model has been proposed that can be applied to all types of connections load-slip. • The BFM can predict the initial stiffness and yield strength of connections quite accurately. Acknowledgement This project was funded by ARC Discovery Grant DP160104092. The support of the ARC to the project is acknowledged with thanks.
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Fragiacomo, A component approach for the hysteretic behaviour of connections in cross-laminated wooden structures, Earthq. Eng. Struct. Dyn. 42 (2013) 2023–2042. [13] S. Pei, J.W. Van De Lindt, M. Popovski, J.W. Berman, J.D. Dolan, J. Ricles, R. Sause, H. Blomgren, D.R. Rammer, Cross-laminated timber for seismic regions: progress and challenges for research and implementation, J. Struct. Eng. (2014), E2514001. [14] B. Dujic, K. Strus, R. Zarnic, A. Ceccotti, Prediction of Dynamic Response of a 7-Storey Massive X-Lam Wooden Building Tested on a Shaking Table, World Conference on Timber…, 2010. [15] M. Okabe, M. Yasumura, K. Kobayashi, T. Haramiishi, Y. Nakashima, K. Fujita, Effect of Vertical Load Under Cyclic Lateral Load Test for Evaluating Sugi CLT Wall Panel, WCTE 2012 - World Conference on Timber Engineering, Auckland, 2012. [16] V. Hristovski, B. Dujic, M. Stojmanovska, V. Mircevska, Full-scale shaking-table tests of X-Lam panel systems and numerical verification — Part 1: specimen 1, J. Struct. Eng. (2012), 12101313044400. [17] M. He, Z. Li, F. Lam, R. Ma, Z. Ma, Experimental investigation on lateral performance of timber-steel hybrid shear wall systems, J. Struct. Eng. 140 (2014) 4014029. [18] M. Fragiacomo, B. Dujic, I. Sustersic, Elastic and ductile design of multi-storey crosslam massive wooden buildings under seismic actions, Eng. Struct. 33 (11) (2011) 3043–3053. [19] C. Loss, M. Piazza, R. Zandonini, Experimental Tests of Cross-Laminated Timber Floors to be Used in Timber-Steel Hybrid Structures, WCTE 2014, Quebec City, Canada, 2014. [20] N. Jacquier, Development and Evaluation of Mechanical Joints for Composite Floor Elements with Cross Laminated Timber. Department of Civil, Mining and Environmental Engineering, Luleå University of Technology, Sweden, 2015. [21] N. Jacquier, U.A. Girhammar, Evaluation of bending tests on composite glulam–CLT beams connected with double-sided punched metal plates and inclined screws, Constr. Build. Mater. 95 (2015) 762–773. [22] BS EN 338, Structural timber, Strength Classes, British Standards Institution (BSI), 2009. [23] ASTM D2555, Standard Practice for Establishing Clear Wood Strength Values, American Society of Testing Materials (ASTM), West Conshohocken, Pennsylvania, United States, 2015. [24] ASTM D3737, Standard Practice for Establishing Allowable Properties for Structural Glued Laminated Timber (Glulam), American Society of Testing Materials (ASTM), West Conshohocken, Pennsylvania, United States, 2012. [25] A. Mårtensson, Mechano-sorptive effects in wooden material, Wood Sci. Technol. 28 (6) (1994) 437–449. [26] F. Dubois, J.M. Husson, N. Sauvat, N. Manfoumbi, Modeling of the viscoelastic mechano-sorptive behavior in wood, Mech. Time-Depend. Mater. 16 (4) (2012) 439–460. [27] AS/NZS 3679.1, Hot-rolled Bars and Sections, 2010. [28] AS/NZS 1393, Coach Screws-Metric Series With ISO Hexagon Heads, 1996. [29] AS 1110.1, ISO Metric Hexagon Bolts and Screws-Product Grades A and B, 2000. [30] AS 1112.1, ISO Metric Hexagon Nuts - Style 1 - Product Grades A and B, 2000. [31] N. Khorsandnia, H.R. Valipour, K. Crews, Experimental and analytical investigation of short-term behaviour of LVL–concrete composite connections and beams, Constr. Build. Mater. 37 (2012) 229–238. [32] BS EN 26891, Timber Structures – Joints Made With Mechanical Fasteners – General Principles for the Determination of Strength and Deformation Characteristics, 1991. [33] B.-H. Xu, A. Bouchaïr, P. Racher, Mechanical behavior and modeling of dowelled steel-to-timber moment-resisting connections, J. Struct. Eng. 141 (2015) 04014165. [34] P. Aune, M. 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Contents lists available at ScienceDirect
Journal of Constructional Steel Research
Load-slip behaviour of steel-cross laminated timber (CLT) composite connections A. Hassanieh, H.R. Valipour ⁎, M.A. Bradford Centre for Infrastructure Engineering and Safety School of Civil and Environmental Engineering, UNSW Australia, Sydney, NSW 2052, Australia
a r t i c l e
i n f o
Article history: Received 1 December 2015 Received in revised form 3 March 2016 Accepted 9 March 2016 Available online xxxx Keywords: Cross laminated timber (CLT) Load-slip Shear-connectors Steel-timber composite (STC)
a b s t r a c t Connecting cross laminated timber (CLT) panels to steel girders using mechanical connectors (e.g. screws and bolts) is an attractive and novel method for developing a fully prefabricated and sustainable hybrid steeltimber composite (STC) floor that can also facilitate future dismantling and recycling of its structural components. Since the structural performance of composite floors (including STC systems) is significantly influenced by strength and stiffness of slab-to-girder composite joints, this study characterises the load-slip behaviour and failure modes of steel-CLT timber composite joints by conducting push-out tests on three different types of STC connections with high-strength bolts, coach screws or a combination of glue and coach screws. Furthermore, the effect of reinforcing the CLT slabs by using steel nail plates on the strength and stiffness of STC joints is evaluated. Empirical models for capturing load-slip behaviour of steel-CLT composite joints with dowel (i.e. screw and bolt) connectors are derived from non-linear regression of experimental data and simple formulae for the strength and stiffness of steel-CLT composite joints with dowel connectors are proposed. Moreover, 1D beam on inelastic foundation finite element (FE) models of the dowel STC connectors are developed and validated against experimental results. © 2016 Elsevier Ltd. All rights reserved.
1. Introduction From a structural engineering perspective, replacing reinforced concrete (RC) slabs with timber flooring can reduce the self-weight of a building significantly, leading to a reduction in energy-intensive concrete foundations as well as to less-demanding rigging and craneage requirements. In addition, lowering the self-weight of the structure significantly reduces the inertial forces induced by seismic actions, particularly on medium to high-rise buildings, and it also facilitates the construction of buildings on soft and/or problematic soils. In steeltimber composite (STC) floors, bolts and self-tapping screws can be utilised to transfer the interface shear between the steel joist and the prefabricated timber slab. STC composite joints with mechanical connectors (i.e. screws and bolts) can also facilitate the dismantling, recycling and reuse of structural components. Cross laminated timber (CLT) is a widely-used innovative building material fabricated by bonding together timber boards to produce a solid panel, with each layer of the panel alternating between longitudinal and transverse lamellae. Owing to the alternation of the grain directions, CLT panels have improved dimensional stability and a much lower variability of their mechanical properties when compared with solid timber. Furthermore, CLT panels have relatively high strength and
⁎ Corresponding author. E-mail address: [email protected] (H.R. Valipour).
stiffness properties in both the longitudinal and transverse directions that make possible the two-way spanning action needed to resist transverse loading. Accordingly, CLT panels are an attractive candidate for developing hybrid STC floors/beams. The structural performance and behaviour of steel-CLT composite floors is significantly influenced by mechanical properties of the CLT panels and the strength and stiffness of composite joints. Several studies have been conducted hitherto to characterise the mechanical properties (tensile, compressive, bending and embedding strengths) of CLT. For example, embedding tests have been performed on CLT panels according to the specifications of EN383-2007 [1], and the results have been used to determine the load carrying capacity of CLT joints [2]. Furthermore, the behaviour and mechanical properties of CLT under compressive loading perpendicular to the grain direction and the effect of boundary line loading have been investigated by Serrano and Enquist [3], and it was shown that the brittle compressive failure mode of CLT panels depends on the loading arrangement. Apart from standardised tension, compression and embedding tests, plate-bending tests on three- and five-layered CLT panels have been carried out to determine the global and local failure mechanisms and the evolution of different crack modes in the CLT lamellae [4]. Based on the plate bending tests, the ductile structural behaviour of CLT plates was also described [4]. In addition to plate bending tests, an advanced laminated plate theory was employed by Stürzenbecher et al. [5] to accurately capture the mechanical behaviour of CLT panels under distributed and concentrated loading.
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Over the past decade, several experimental and theoretical studies have focused on the behaviour of CLT panel-to-panel connections. The load carrying capacity and ductility of double lap CLT joints with dowel connectors have been studied by Blaß et al. [6]. Large embedding deformations were observed in the CLT lamellae, and using punched metal plate fasteners or glued-on boards was found to be effective in improving the embedding strength of the CLT panels as well as the load carrying capacity of double lapped CLT joints [6]. Moreover, monotonic and cyclic shear tests on lapped and orthogonally connected CLT/X-Lam panels (representing walls or floors) with screws and steel angle brackets were performed by Gavric et al. [7,8] to evaluate the strength, stiffness, energy dissipation and ductility of CLT/X-Lam wall-to-wall, wall-to-floor/foundation and floor-to-floor connections. Similarly, the cyclic/seismic behaviour of CLT shear wall-to-floor connections with steel brackets and mechanical fasteners in both parallel and perpendicular to the grain directions has been investigated experimentally and theoretically by Shen et al. [9] and Schneider et al. [10]. The results of cyclic tests conducted by different researchers have demonstrated the reasonable seismic performance of CLT shear walls with mechanical connectors such as nails and screws [10,11]. In addition to behaviour of CLT panel-to-panel connections, several studies have been performed experimentally and numerically on the seismic performance of CLT structures, and the influence of connections' behaviour on the global response of the buildings with CLT shear walls and floors [8,12–18]. However, experimental data on behaviour of steel-timber and concrete-timber composite floors with CLT slabs and dowel connectors is very limited. Very recently, Loss et al. [19] performed symmetric push-out tests to characterise the behaviour of different connections in steel-CLT composite floors [19]. Furthermore, the load-slip behaviour of timber-timber and concrete-timber composite connections with CLT slabs have been investigated by conducting push-out tests on asymmetric specimens [20,21]. This paper investigates the short-term load-slip, peak load carrying capacity, stiffness and failure mode of steel-CLT composite joints by conducting symmetric push-out tests. In the specimens tested, the CLT panels are connected to the flanges of the steel profile using bolts, coach screws and/or glue. Non-linear regression of the experimental results is undertaken to derive empirical load-slip formulae for steel-CLT composite joints with coach screws and high-strength bolts. The empirical load-slip models developed can be incorporated in finite element models and used for the nonlinear analysis of steel-CLT composite beams/floors with dowel connectors. In addition, 1D beam on inelastic foundation finite element (FE) models of the screw connectors are developed and validated against experimental push-out tests.
2. Experimental program 2.1. Specimen details The push-out tests were conducted on nine different groups of symmetric steel-CLT specimens that comprise of two CLT panels connected to the flanges of a steel profile (Fig. 1). The specimens were categorised with respect to the type of connectors (i.e. bolts and screws with and without glue) and application of nail plate fasteners for reinforcing the CLT panels. Furthermore, an end connection was designed to be used at the two ends of the simply supported composite beams. The end connection comprised of a 50 × 50 × 5 mm steel box connected to the flange of the steel girder by two high-strength friction grip Grade 8.8 bolts that were prestressed to 60% of their yield load (Fig. 1). The geometry, details of the specimens, test set up, location of loads and supports and type of dowel connectors are shown in Fig. 1 and in Table 1. The primary variables within the different push-out specimens were the
Fig. 1. Configuration and cross-sections of nine different steel-CLT composite joints for push-out tests.
diameter and type of the mechanical connectors (i.e. bolts or screws) and the use of glue or punched steel plate fasteners in conjunction with screws.
Table 1 Size and configuration of connectors and loading direction in the push-out specimens. Group
Mechanical connectors
With nail plate Without nail pate Glued End connection
Screw 12, 16, 20 & bolt 12, 16 Screw 12, 16 Glue + screw 16 50 × 50 × 5 box connection + bolt 12 + screw 12
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2.2. Material properties 2.2.1. Cross laminated timber (CLT) The CLT panels were 120 mm thick and they were made from five layers (Fig. 2a) of Spruce lamellae of strength class C24 according to BS EN338 [22]. The orientation and thickness of each lamella are shown in Fig. 2a. The CLT panels were 400 mm wide and 600 mm long and in the push-out tests the CLT panels were loaded in the direction parallel to the grain of the first layer (Fig. 1). The CLT panels had an average moisture content (MC) of 12% (±2%) and density of 490 kg/m3 at the time of fabrication and testing of the STC specimens. The nominal
mechanical properties of the CLT panels (at 12% MC) are given in Table 2. The mechanical properties of timber/engineered wood in different standards such as ASTM D2555 [23] and ASTM D3737 [24] are typically given as ratios of mechanical properties at a specific MC, i.e. 12%. It is well-established that mechanical properties and behaviour of timber and engineered wood products can be influenced by the moisture content and its variations [25,26]. Generally, the strength and elastic modulus of timber decline as the moisture content of timber increases. This change in timber modulus of elasticity and/or strength can affect the load-slip and peak load capacity of the STC system. 2.2.2. Steel profile The push-out specimens were fabricated by connecting CLT panels to Australian 310UB32 hot rolled I-sections that comply with the specifications of AS/NZS 3679.1 [27]. The dimensions, yield strength and ultimate strength of the 310UB32 cross-section are given in Table 3. 2.2.3. Dowel connectors Hexagonal high-strength Grade 8.8 bolts having diameters of 12 mm and 16 mm, and hexagonal coach screws having diameters of 12 mm, 16 mm and 20 mm were used in the fabrication of the specimens (Fig. 2b). The screws were made of Grade 4.6 steel with a yield strength of 240 MPa and an ultimate tensile strength of 400 MPa, and they comply with the minimum requirements of AS/NZS 1393 [28]. The highstrength bolts had a proof yield strength of 660 MPa and an ultimate tensile strength of 830 MPa, as per the specifications of AS1110.1 [29] and AS1112.1 [30]. The coach screws were 100 mm long and the bolts were 110 mm long. 2.2.4. Nail plate The specimens with 12 and 16 mm screw connectors were tested twice, i.e. with and without nail plates. The nail plates were used to reinforce the CLT panels around the connector hole and to improve the stiffness and strength of the STC joints. The nail plates were 100 × 50 × 1 mm (Fig. 2c) and made of Grade G300 steel with a characteristic yield strength of 300 MPa and a tensile strength of 340 MPa. In the composite joints with 12 and 16 mm screws, one and two nail plates per screw were used respectively to reinforce the CLT panels. The nail plates were installed on the CLT panels simply by hammering. After installation of the nail plates, the holes required for installation of dowel fasteners (i.e. bolts or screws) were drilled in the nail plates and then the CLT panels were mounted on steel beam flange and the coach screws/bolts were installed. Only one nail plate was used in conjunction with each 12 mm shear connector, however, two nail plates installed in a cruciform configuration were used in conjunction with 16 mm dowel connectors. 2.2.5. Glue In some of the specimens, a non-sag gel type epoxy was used in conjunction with 16 mm coach screws to provide full composite action between the CLT panels and steel profile. The mechanical properties and curing requirements of the epoxy are provided in Table 4. 2.3. Test setup and instrumentation of the specimens
Fig. 2. (a) 5-Layer cross laminated timber (CLT) panel (b) bolt and coach screw connectors and (c) nail plate.
A symmetric CLT-steel-CLT arrangement (Fig. 1) for the fabrication of specimens and test set up was adopted in this study to prevent unsymmetrical load distribution on the sample and to minimise unwanted friction between different parts of the specimen [31]. The symmetric configuration of the specimens can provide a close to uniform distribution of load with predictable friction effects on the load-slip response, and it can also facilitate fabrication of the push-out test specimens. The relative displacement (or slip) between the CLT panels and steel profile was measured using four linear variable differential transformers (LVDTs) having 100 mm maximum stroke. Furthermore, an LVDT was used to measure the vertical movement of testing machine actuator as
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Table 2 Mechanical properties of CLT (in MPa). Bending (fb)
Tension parallel to grain (ft)
Shear in beams (fs)
Compression parallel to grain (fc)
Compression perpendicular to grain (fp)
Elastic modulus (E)
Modulus of rigidity (G)
24
16.5
4.6
24
2.7
12,000
690
Table 3 Mechanical and geometrical properties of steel profile. Section depth d (mm)
Flange width bf (mm)
Flange thickness tf (mm)
Web thickness tw (mm)
Yield strength fy (MPa)
Ultimate strength fu (MPa)
Elastic modulus Es (MPa)
298 mm
149 mm
8 mm
5.5 mm
320 MPa
450 MPa
205 GPa
well as the specimens. A steel-CLT composite specimen placed in the testing rig and location of LVDTs mounted on the specimen are shown in Fig. 3. One of the main objectives of this study was to develop a sustainable and demountable steel-CLT composite floor with respect to practical construction considerations. Accordingly, the CLT panels and flanges of the steel profiles were predrilled and the specimens were then assembled following a procedure similar to that representative of current construction practice. The dimeter of the holes predrilled in the flanges of the steel profile was 0.2 mm larger than the dimeter of the screws/ bolts. The diameter of the holes predrilled in the CLT panels was 2 mm smaller than the diameter of the screws, and 1 mm to 2 mm larger than the diameter of the bolts. In the bolted connections, a 50 mm × 50 mm washer was used to prevent crushing of the CLT in the direction perpendicular to the plane of the timber lamellae. A calibrated torque wrench was used to tighten the high-strength bolts to induce a post-tensioning force of 10 kN equivalent to 0.14fy in the 12 mm bolts and 0.08fy in the 16 mm bolts (fy being the yield stress of the bolt). The post-tensioning force was calculated based on a design crushing strength of approximately 4 MPa for CLT and area of 2500 mm2 for the square washers. In the glued steel-CLT composite joints, the epoxy adhesive was applied over a 450 mm by 152 mm area on the surface of the CLT panels and after mounting the CLT panels, the specimen was left for one hour and the 16 mm screws were then installed. The thickness of the adhesive layer after drying was around 2 mm.
3. Discussion of test results 3.1. Modes of failure Four distinctive modes of failure, identified as Modes I, II, III and IV, were observed in the push-out tests on timber-timber, timberconcrete and timber-steel joints with dowel connectors [33–36]. These modes of failure have been also recognised in timber structures standard BS EN 26891:1991 [32]. The occurrence of each mode of failure depends on the slenderness of the connectors represented by L/D (with L being the length and D the diameter of the connector) and the strength ratio fs/bs (with fs being material strength of the connector and bs the bearing strength of the primary element [35]). Failure Mode I is associated with crushing of timber in the vicinity of the dowel connectors without significant deformation in the dowels (Fig. 5). In failure Mode II, a plastic hinge forms within the dowel and crushing of timber occurs close to the surface of the timber element, whilst in Mode III, crushing of the timber is associated with the formation of two plastic hinges, one at the middle and the other one close to the end of the connector (Fig. 5). Failure mode IV is associated with fracture of the connectors and significant crushing of the timber. It is noteworthy that failure Modes I and IV are brittle, however, failure Modes II and III can be ductile or semi-ductile depending on L/D and fs/bs. The dominant failure mode of the steel-CLT composite joints with 12 mm screws (without a reinforcing nail plate) that were tested was Mode III, while in the composite joints with 16 mm and 20 mm screws (without a reinforcing nail plate), the failure mode was Mode II (Fig. 6a– c). In steel-CLT composite joints with bolted connectors (including the
2.4. Loading procedure The specimens were loaded following a procedure specified in the Eurocode for timber structures [32]. In the first stage, the load was increased from 0 to 0.4Fu (with Fu being the estimated ultimate load) over 120 s and the load was maintained at 0.4Fu for 30 s (Fig. 4). The second stage involved unloading of the specimen from 0.4Fu to 0.1Fu and then maintaining the load at 0.1Fu for 30 s. In the final stage, testing was conducted under a load control regime up to 0.7Fu, and then reverted to displacement-control (with an average displacement rate of 2 mm/s) and continued until failure of the specimen (Fig. 4). The loading procedure in each test was redefined with respect to the ultimate load Fu observed in the previous test.
Table 4 Mechanical properties and curing requirements of the epoxy adhesive. Tensile strength
Compressive strength
Shear strength
Coefficient of linear expansion
Minimum cure time at 25 °C
30 MPa
70 MPa
15 MPa
60 × 10−6 mm °C
24 h
Fig. 3. Push-out test setup and instrumentation.
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Fig. 4. BS EN 26891 loading procedure adopted for the push-out tests [32].
joints with end 50 × 50 × 5 mm steel box connections), Mode IV was the dominant failure mode (Fig. 6d). Using nail plates to reinforce the CLT panels did not alter the failure mode of the joins with screw connectors, but the nail plates reduced the ductility of joints slightly. The reduction in the ductility of joints with nail plates was evident from the sudden drop of the load after the peak load was attained. The failure mode of joints with reinforcing nail plates was associated with rupture of nail plate, localised crushing of the timber and the formation of one or two plastic hinges in the screws, depending on the dimeter of the screws (Fig. 7a). In general, the joints with screw connectors exhibited a relatively ductile behaviour, whereas the joints with bolted connectors or combination of glue and screws produced a brittle mode of failure that was associated with fracture of the bolts or glue (see Fig. 7b) and separation of the CLT panels from steel profile. It is noteworthy that the steel profiles and the bolt or screw holes in the flange of steel profile did not exhibit any sign of deformation during or after the test. 3.2. Load-slip response of connections It is well known that the variability of timer mechanical properties and the method of fabrication can significantly affect the structural behaviour of timber connections. Accordingly, three identical specimens
Fig. 5. Characteristics of failure modes I, II and III observed in the steel-CLT composite joints.
Fig. 6. Failure mode of steel-CLT composite joints with (a) screw S12, (b) screw S16 and (c) bolt B12 and (d) bolt B16 connectors.
A. Hassanieh et al. / Journal of Constructional Steel Research 122 (2016) 110–121
Fig. 7. (a) Crushing of timber and nail plate rupture in the steel-CLT composite joints with screw connectors and reinforcing nail plate and (b) fracture in the glued joints.
Fig. 8. Load-slip response of steel-CLT composite joint with (a) 2 × S12 (12 mm screw) (b) 2 × S16 (16 mm screw) and (c) 2 × S20 (20 mm screw).
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were fabricated for each type of steel-CLT composite joint (Table 1 and Fig. 1), and the specimens were tested under identical loading conditions and test set up, to ensure the precision and reproducibility of the experimental data (i.e. load-slip response, stiffness under service load condition, peak load capacity and failure mode). The load-slip responses of the STC joints with screws (with and without a reinforcing nail plate) and bolted connectors obtained from the three identical specimens are shown in Figs. 8 to 10 respectively. Furthermore, the mean load-slip curves are also shown in Figs. 8 to 10, which demonstrate clearly the small variation in the peak load capacity and of the load-slip response of the steel-CLT composite joints. In the bolted steel-CLT composite joints, the friction between the steel profile and the CLT panel mobilised by the post-tensioning of the high strength bolts has provided full composite action at the early stages of the load-slip evolution, where no slip occurs up to a shear force of 11 kN (Fig. 9). The load corresponding to first slip (see Fig. 9) in the bolted STC joints depends on the magnitude of post-tensioning force induced in the high-strength bolts. Accordingly, increasing the initial post-tensioning force in the bolted connectors can increase the initial stiffness (initial slope of the load-slip behaviour) of the bolted connectors, however, the magnitude of post-tensioning force has minor influence on the peak load carrying capacity of the bolted STC joints. It is noteworthy that the magnitude of post-tensioning force in the bolts must be specified/limited with respect to the compressive strength of CLT panels to ensure that crushing/punching (e.g. Fig. 6d) of CLT in the direction perpendicular to the plane of lamella would not happen.
Fig. 9. Load-slip response of steel-CLT composite joint with (a) 2 × B12 (12 mm bolt) and (b) 2 × B16 (16 mm bolt).
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Fig. 10. Load-slip response of steel-CLT composite joint with reinforcing nail plates and (a) 2 × S12 (12 mm screw) and (b) 2 × S16 (16 mm screw) connectors.
Apart from steel-CLT joints with dowel (i.e. screw and bolt) connectors, a composite joint with a combination of glue and screws and also a composite joint with friction grip end 50 × 50 × 5 mm steel box connections were fabricated and tested. The experimental load-slip and shear stress-slip results for the glued joints and the experimental load-slip response of the composite joints with friction grip end 50 × 50 × 5 mm steel box connections are shown in Figs. 11 and 12, respectively. Moreover, the mean and coefficient of variation (CoV) of the peak load carrying capacity and the mean of the serviceability stiffness (represented by the slope of the load-slip curve between 10% and 60% of the peak load capacity, i.e. the ks,0.6 slip moduli) for the steel-CLT composite joints are given in Table 5. The maximum CoV of the peak load carrying capacity of the steel-CLT composite joints with screw and bolt connectors was limited to CoVpeak = 8.8% and 3.9%, respectively. With regard to the load-slip response and peak load capacity of the tested composite joints (Figs. 8 to 12 and Table 5), the following conclusions can be established: • The glued steel-CLT composite joints exhibited significantly higher strengths and stiffnesses compared to the joints with only mechanical connectors (i.e. dowelled or friction grip end box connections). Furthermore, the friction grip end box connections can have higher peak load carrying capacity and stiffness than conventional dowelled connectors. • The load carrying capacity of steel-CLT joints with bolted connectors is higher than for joints with screws, because of the higher strength of the steel used for manufacturing the Grade 8.8 bolts compared to the mild steel used for manufacturing the screws. • The experimental load-slip responses of the joints with screw connectors (Figs. 8 and 10) exhibited close to elastic-perfectly plastic behaviour, whilst the steel-CLT joints with bolted connectors showed some hardening characteristics (see Fig. 9).
Fig. 11. (a) Load-slip and (b) stress-slip response of glued steel-CLT composite joints.
• Reinforcing the CLT panels by nail plates can enhance the load carrying capacity and particularly the serviceability stiffnesses of steelCLT composite joints with screws. A minimum strength enhancement of 28% was achieved in composite joints with S16 screws where two nail plates per S16 screw hole were used to reinforce the CLT panel. Furthermore, a minimum stiffness increase of 70% for S12 and 226% for S16 connectors was observed. However, reinforcing the CLT panels by nail plates can reduce the ductility of steel-CLT composite joints. The less ductile behaviour of joints with nail plates is evident from the post-peak softening of load-slip response (see Fig. 10). 4. Analytical load-slip model for steel-CLT composite joints with dowel connectors In this section, empirical load-slip models for the tested steel-CLT composite joints with dowel connectors are developed by non-linear
Fig. 12. Load-slip response of the steel-CLT composite joints with end 50 × 50 × 5 mm steel box connection.
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Table 5 Mean and CoV (in %) of the peak load capacity and mean of the slip moduli ks,0.6a for the STC joints. Screwb
Nail plate Slip moduli ks,0.6 (kN/mm) Peak load capacity (kN)
Without With Without With
Boltb
S12
S16
S20
B12
B16
9.52 16.23 35.78 [3.7] 38.88 [6.3]
8.21 26.8 49.87 [8.8] 63.98 [3.8]
6.11 – 67.52 [2.4]
6.26 – 79.88 [2.2]
9.66 – 123.4 [3.9]
Box 50 × 5 × 5 + 12 mm bolt 34.94 – 131.13 –
a The slip moduli ks, 0.6 is used as a representative of the connection stiffness within the service range of loading. The ks, 0.6 slip moduli is the slope of the load-slip curve between 10% and 60% of the peak load capacity. b The values in [] are CoV (in %) obtained from test results on three identical specimens.
regression analysis of the push-out test results. The proposed load-slip function is inspired by Ramberg-Osgood [37] model that has been widely used in analysis of structural steel components [38]. Essentially, a post-peak softening branch has been incorporated into the RambergOsgood model to represent both hardening and softening behaviour of steel-timber composite connections with just one mathematical function. The proposed model can be easily incorporated into componentbased finite element models and used for the non-linear analysis of hybrid steel-CLT composite beams/floors with coach screw and bolt connectors. For the sake of ease of computer implementation, a seven-parameter model built on three asymptotic lines (see Fig. 13) is proposed, being given by k0 −kp s kp þ ks s þ f ¼n 1 h h i n2 on1 −ks s; i n1 on1 n 2 1 þ k0 −kp fs 1 þ kp þ ks f −s f 0
1
ð1Þ
0
where f is the shear force, s the slip, k0 the initial stiffness, kp the prepeak stiffness, ks the post-peak stiffness, f0 the first reference shear force corresponding to the pre-peak branch and f1 the second reference shear force corresponding to the post-peak branch, and where n1 and n2 are two parameters that control the curvature of the first and second curves, respectively (Fig. 13). It is noteworthy that the stiffness values
k0, kp and ks can be positive, negative or zero and accordingly the proposed empirical model can capture the pre-peak as well as the postpeak load-slip response of the steel-CLT composite joints with dowel connectors. The non-linear regression based on the least squares method with trust-region algorithm and bisquare robustness available in MATLAB software was employed to fit the proposed function to the experimental data and determine the values of seven different input parameters, i.e. k0, kp, ks, f0, f1, n1 and n2. In the first step, the mean experimental loadslip for each group of steel-CLT joints was obtained. In the second step, the initial values and upper and lower bounds for the seven unknowns are defined with respect to the mean of experimental loadslip and the values of each parameter within a 95% confidence bound is calculated using the adopted non-linear regression procedure. The seven input parameters required for establishing the analytical loadslip model of the steel-CLT composite joints with mechanical connectors (i.e. screw, bolt and end box friction grip connection) are given in Table 6. In addition, the R-square values for the calibrated load-slip curves are provided in Table 6. It is seen that the R-square values for all calibrated models are greater than 0.9798, demonstrating goodness-of-fit and close correlation between the analytical model and experimental data. The excellent correlation between the proposed analytical models and the experimental results is also evident from the load-slip curves shown in Fig. 14. In addition, a linear regression was carried out on the experimental slip moduli ks,0.6 and peak load carrying capacity of the joints with coach screw connectors versus the diameter of screws and the results are shown in Fig. 15. Accordingly, a formula is proposed for calculating the ultimate load carrying capacity Pu as 2
P u ¼ 0:123d þ 18:28
ð2Þ
and the serviceability stiffness (represented by slip moduli ks,0.6) of the steel-CLT composite joints with screw connectors as 2
ks;0:6 ¼ 3:94d −11:96
Fig. 13. Proposed analytical model for load-slip behaviour of steel-timber composite joints.
ð3Þ
where, Pu is the ultimate strength of peak load carrying capacity (in kN), d the screw dimeter (in mm) and ks,0.6 the slip modulus (in kN/mm).
Table 6 Input parameters for the analytical load-slip model of the steel-CLT composite joints with mechanical connectors (i.e. screw, bolted and end box friction grip connection). Connector
f0 f1 k0 kp ks n1 n2 R2
Screw without nail plate
Screw with nail plate
Box 50 × 50 × 5 + 12 mm bolt
S12
S16
S20
B12
Bolt B16
S12
S16
B12
26.7 42.4 47.7 0.741 0.337 1.03 8.3 0.9985
49.5 284.5 198.8 1.569 0.778 0.5 0.52 0.9943
52.4 73.8 310.7 1.836 ‐0.069 0.45 9.1 0.9998
6.8 40.9 114.5 11.3 -1.396 8.48 2.75 0.9987
5.5 69.8 122.1 16.27 -1.097 2.24 3.35 0.9998
38.2 85.5 41.5 0.144 1.654 1.19 4.015 0.9997
37.4 65.5 64.5 52.8 0.553 6.294 2.13 0.9798
11.1 301 372 52 9.6 5 1.14 0.998
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Fig. 15. Correlation between coach screw diameter and mean of experimental (a) peak load capacity and (b) slip moduli ks,0.6.
(Fig. 16a). The screws are modelled by a non-linear force-based beamcolumn element with distributed plasticity which is available in the OpenSEES [45] library. The foundation behaviour is modelled by zerolength elements that connect the nodes on the beam (connector) to fixed supports by springs with predefined force-displacement relationships (Fig. 16a). For modelling friction between the screw and timber along the axis of screw, another zero-length element was used. The behaviour of the springs used for capturing the friction effect was obtained
Fig. 14. Correlation between analytical model and mean of experimental load-slip response for steel-CLT composite joints with (a) screw connectors, (b) bolted connections and (c) screw connector and reinforcing nail plate.
5. Modelling the connectors using beam on (in)elastic foundation theory The theory of beams on (in) elastic/inelastic foundation can be used to model and predict the behaviour of shear connectors in concrete or timber [35,39,40]. In this modelling strategy, the shear connector (i.e. bolt, screw and nail) is considered as a beam placed on a foundation made of concrete or timber. Accordingly, several studies have used 1D beam elements in conjunction with linear/non-linear springs (representing the foundation) to capture the behaviour of shear connectors in timber-concrete, timber-timber and steel-timber composite joints [41–44]. The main input parameters required for developing such models are the yield strength of the connectors and the embedding stiffness and strength of timber that can be obtained from embedding test results [43]. In this study, the Beam-on-Foundation model (BFM) is employed to capture the behaviour of steel-CLT composite joints with coach screws
Fig. 16. (a) BFM model (b) deformed connection.
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Fig. 17. Adopted uniaxial stress-strain relationship for steel.
from withdrawal tests [2] and the withdrawal capacity R (in N) of the axially loaded screws was obtained from 0:8 0:9 lef
ð4Þ
R ¼ 31 d
where d is the screw diameter (in mm) and lefis effective penetration length (in mm). The effective penetration length lef is obtained by subtracting the length of the cone shaped part of the screw shank (approximately 5 mm) and the steel profile flange thickness (8 mm) from the total length of the screw. It is noteworthy that the effective penetration length can have some influence on the load-slip behaviour and peak load capacity of the screw connectors that requires further investigation. In the FE model, it is assumed that the steel profile flange can provide a rotationally fixed support for the connectors (Fig. 16), owing to its significantly larger flexural stiffness compared to the flexural stiffness of the connectors (screws). To capture the yielding of steel, the screw (beam) cross-sections were discretised to layers/fibres and a uniaxial constitutive law was assigned to each layer/fibre. A multilinear stress-strain relationship (see Fig. 17) was used for modelling Grade 4.6 steel. Moreover, the behaviour of timber foundation was characterised using the results of embedding test performed by Nakashima et al. [46] on CLT specimens. The embedding strength versus displacement curves in Nakashima et al. [46] were converted to force-displacement curves for foundation springs in both parallel and perpendicular to the grain direction. The wood foundation parameters and embedding stress versus displacement curves adopted for developing the force-displacement relationship of the foundation springs are provided in Table 7 and Fig. 18, respectively. It is noteworthy that the layer-wise characteristics of CLT panels were considered in the beam-on-foundation finite element models by assigning force-displacement relationships in parallel and perpendicular to the grain directions to different parts of the connector as shown in Fig. 16. The load-slip response of the steel-CLT composite joints with screw connectors predicted by the beam on foundation FE models are compared with the experimental results in Fig. 19 and a reasonably good correlation between the FE model predictions and experimental data (particularly the initial stiffness and peak load capacity) is observable. However, the yield load of the composite joints predicted by the FE Table 7 Adopted wood foundation parameters. f0l MPa
f0y MPa
f90l MPa
f90y MPa
d0l mm
d0y mm
d90l mm
d90y mm
α0
α90
21
33
9
12
0.6
1.31
0.78
1.21
0.01
0.05
Fig. 18. Adopted wood foundation stress-displacement behaviour for (a) parallel and (b) perpendicular to the grain direction.
model has only a mediocre accuracy. The discrepancy between the FE and experimental results can be attributed to variability of steel and timber mechanical properties with respect to diameter of screws (size effect) and also the influence of drilling and installation methods on the mechanical properties of timber in the vicinity of the connector holes. 6. Conclusions The study illustrated the general behaviours of various connection types in CLT-to-Steel composite parts as the most critical element in timber-steel hybrid structures. The main points of the study are the following: • The load-slip response of coach screw connections demonstrated ductile behaviour, and typically failure occurs after a large post-peak branch in the load-slip curve. Local timber crushing and developing plastic hinges in coach screws causes failure of whole specimen. The pre-stressed bolted connections show a ductile response and failure is associated with rupture and sudden drop in load slip response. • Glued connections are much stiffer and stronger than screw and bolted connections, but they have a brittle failure that undesirable. Using glued connections in conjunction with the coach screws leads to more ductile behaviour of the connection.
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References
Fig. 19. Comparison of load-slip response predicted by beam on foundation FE models for (a) 2 × S12, (b) 2 × S16 and (c) 2 × S20 connectors.
• Reinforcing the coach screws connections with nail plates can increase the initial stiffness and strength of the connections. • The end connection demonstrated considerably higher stiffness and strength. This connection can be used in beam/floors where positive bending moment is applied on these elements. Although it is considered as the first and last shear connections in composite beams, it may be used in the first and last spots of each panel segment with some considerations. • The analytical model has been proposed that can be applied to all types of connections load-slip. • The BFM can predict the initial stiffness and yield strength of connections quite accurately. Acknowledgement This project was funded by ARC Discovery Grant DP160104092. The support of the ARC to the project is acknowledged with thanks.
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