Mechanical performance of novel steel one-sided bolted joints in shear

Journal of Constructional Steel Research 165 (2020) 105815

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Mechanical performance of novel steel one-sided bolted joints in shear Chuqi Wan a, Yu Bai a, *, Chenting Ding a, Lei Zhu b, Lu Yang c a

Department of Civil Engineering, Monash University, Clayton, VIC3800, Australia School of Civil and Transportation Engineering, Beijing University of Civil Engineering and Architecture, Beijing, China c College of Architecture and Civil Engineering, Beijing University of Technology, Beijing, China b

a r t i c l e i n f o

a b s t r a c t

Article history: Received 1 September 2019 Received in revised form 17 October 2019 Accepted 18 October 2019 Available online xxx

This paper introduces an innovative one-sided bolt with ellipse head and examines its mechanical performance based on experimental and numerical studies of double lapped joint configuration. The design of one-sided bolt mechanism and geometry was completed through detailed Finite Element (FE) analysis considering contact behaviours and bolt pre-tensioning. Four bolt head geometries were assessed and comparatively studied in terms of yielding capacity, slip distance and stress concentration; and two bolt diameters and two installation configurations were introduced for each bolt head geometry. Experiments were then conducted based on the specimens manufactured with the optimised geometry for validation of the design and installation considerations, with comparison to ordinary bolts. It was found that the optimal aspect ratio of the ellipse bolt head is about 1.7 and orthotropic installation configuration was recommended. Such innovative one-sided bolts with simple manufacturing and installation sequence exhibit comparable shear performance with ordinary bolts, enabling the use of closed sections for construction with easy assembly and disassembly. © 2019 Elsevier Ltd. All rights reserved.

Keywords: One-sided bolt Closed section member Double lapped joint Shear performance Hollow section members

1. Introduction The applications of hollow sections have gained broad recognition in steel structure construction considering both functional and architectural requirements. Being a closed section form, hollow sections exhibit better resistance against torsional and global buckling compared with open sections (i.e. I or channel profiles) [1e4]. The closed shape improves the aesthetic appearance and reduces the area for corrosion protection [5,6]. However, the use of hollow section members presents challenges in connections to other structural members. Welding is usually used in such circumstances; compared to welded connections, bolted connections are regarded as an alternative with easier quality control and onsite applications [7e12]. Conventional bolted connections require go-through bolts across the hollow section members otherwise it becomes inconvenient due to limited access inside the hollow section to tighten the nut. Blind or one-sided bolting systems permit bolt installation and

* Corresponding author. E-mail addresses: [email protected] (C. Wan), [email protected] (Y. Bai), [email protected] (C. Ding), [email protected] (L. Zhu), [email protected] (L. Yang). https://doi.org/10.1016/j.jcsr.2019.105815 0143-974X/© 2019 Elsevier Ltd. All rights reserved.

tightening from one side of the connection. This convenience can be very useful when only one side of the connection is accessible such as in the case of hollow section connections. There are several types of existing one-sided bolting systems such as Flowdrill [13], Huck Blind Bolt [14], Lindapter Hollobolt [15], Blind Bolt [16]and Ajax Oneside [17]. One-sided bolting systems in the present market often require special installation equipment and techniques. Each of them differs from one another significantly, demonstrating varying solutions possible and different requirements for different bolting systems individually. With the limited number of one-sided bolt designs available and the associated complex manufacturing and installation sequences, there is a clear opportunity for development of a new one-sided bolt design with potential cost reduction and easier installation method. Flowdrill is a thermal drilling technique that can achieve bolt installation from one side of the connection, by locally thickening the section around bolt hole between 1.5 and 2 times with tungsten-carbide friction drill and threading the bolt hole on the connected members using roll tap [18e21]. This method allows the use of conventional bolts without tightening nuts. To achieve full tension capacity of grade 8.8 bolts in flowdrilled holes, a minimum tube wall thickness should be permitted to ensure thread stripping failure does not occur prior to bolt failure [22]. These thicknesses

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are greater than those of the hollow section columns normally used in low to medium rise structures [21]. Therefore, the flowdrill process may be only limited to certain hollow sections, for example for infrastructure applications and may not be convenient for onsite practices. Blind oversized mechanically locked bolt (BOM), Huck high strength blind bolt (HSBB) and Ultra-twist are three types of onesided fasteners with similar components, tightening mechanism and installation method. These fasteners achieve one-side assembly by deformable component forming lock head on the inaccessible side of the connection [23,24]. The installation sequences of the BOM fasteners are shown as a representative in Fig. 1a. These three kinds of one-sided fasteners rely on special electric installation wrench on site. Possible premature damage of deformable components under tension limits the tensile capacity [25]. Hollobolt is another kind of one-sided fastener with spreadable legs (see Fig. 1b). The bolt tail is inserted into a predrilled hole and the torque is applied to the bolt through a torque wrench on the bolt head side [26]. The torque applied draws the cone moving toward the exterior side of the connection, expanding the legs which form a fixing against pulling out [27e29]. The reverse mechanism hollobolt (RMH) and extended hollobolt (EHB) are two types of one-sided fasteners evolved from Hollobolt. RMH has similar components as Hollobolt but the expanding part is inverted. Tightening of the bolt head causes flaring of sleeve legs which produces more effective clamping force than that provided by Hollobolt [15]. EHB is featured by extended bolt shank and an anchorage nut in the concrete fill [30]. The strength is improved by failure mode changing from bolt pull-out to bolt shank tensile fracture and the stiffness is enhanced by the increased bolt anchorage and threaded nut at the end for closed section members with concrete infill [11,31,32]. Hollobolt is not reusable as the flaring of sleeve legs deforms permanently. Disassembly of this kind of one-sided bolts can only be achieved by severe damage of the sleeves. Fig. 1c shows a Blind bolt with its locking anchors retracted and released. This type of one-sided bolt permits one side installation by the retractable locking anchor. On installation of the bolt, the locking anchor is retracted and the bolt tail is inserted into the hole on connected members. The locking anchor is then released, allowing the bolt to be tightened [33]. The use of locking anchor compromises the area of the shear plane across slot and this may result in lower shear failure load compared to ordinary bolts [34,35]. There is also a maximum torque for each bolt diameter considering the loading carrying capacity of locking anchor [36]. The clamping thickness is governed by the distance between the locking anchor and thread as shown in Fig. 1c. This will limit the

applications of such bolts as various bolts will be required corresponding to various clamping thicknesses. The Ajax Oneside (Ajax) is another type of one-sided bolts. It compromises several components including bolt, shear sleeve, collapsible washer and nut and needs to be assembled by special installation tool [37]. The main mechanism relies on the special collapsible washer that can be folded before insertion and unfolded after installation with the help of the special tool. The collapsible washer is then bearing against the interior wall with the bolt head and the nut is tightened from outside [32]. The required clearance for Ajax fasteners is larger than that for ordinary bolts, i.e. 18 mm required for M12 Ajax bolts and 13 mm for ordinary M12 bolts. The oversized bolt hole may lead to large localised deformation under loading, causing bolt pull-out failure [38]. Larger values of the minimum edge and pitch distance are required as well in comparison to ordinary bolts. Due to the requirement of the special installation tool, a minimum cavity length is needed for installation [39]. An innovative one-sided bolt without excessive components or special installation techniques is developed and introduced in this work. It presents the detailed design of bolt mechanism and geometry and further experimental and numerical studies for such innovative one-sided bolts, through the double lapped joint configuration in shear. Specimens were designed for numerical modelling to investigate the effect of different bolt head geometry, installation configuration and bolt diameter. An optimal bolt head geometry was then determined through the discussion of yielding capacity, slip distance and stress concentration. Experimental results were further conducted to understand the shear performance of double lapped joints with such innovative one-sided bolts and results were compared and discussed with the Finite Element (FE) results regarding load-deformation behaviour and local strain responses. The effects of installation configuration and loading direction were clarified based on the results from both experimental study and FE analysis. A comparative study with ordinary bolt was finally carried out to indicate the differences in the mechanical responses of the double lapped joints using ordinary bolts or such innovative one-sided bolts. 2. Design of one-sided bolt mechanism and geometry The design concept of the proposed one-sided bolt is to modify the geometry of the bolt head and to ensure bolt locking by the bolt head in the inaccessible side (see Fig. 2). This can be achieved by an ellipse bolt head for example that is inserted into the bolt hole on connected members and then rotated by ninety degrees. This permits the bolt head fixing against the interior wall, due to the

Fig. 1. Installation sequence of: a) BOM (Photo courtesy of AFS Huck, America) b) Hollobolt (Lindapter international, UK) c) Blind bolt (Blindbolt UK).

C. Wan et al. / Journal of Constructional Steel Research 165 (2020) 105815

Fig. 2. Proposed one-sided bolt.

difference between the long and short axis of the ellipse bolt head. Then nut can be tightened from the accessible side. A slot cut (pointed in Fig. 2) directing the long axis of the bolt head will be prepared at the tail of the bolt shank to ensure proper installation. 2.1. Design parameters In order to determine an optimal bolt head design, numerical analysis was first conducted to understand the shear behaviour of such one-sided bolted double lapped joints using FE approach. As proposed, an ellipse bolt head was considered for the innovative one-sided bolt while with the bolt shank and washer consistent to ordinary bolts. For each bolt head, the short axis equals to the bolt shank diameter. Four different ratios between long axis and short axis of the ellipse bolt head (i.e. aspect ratio) were examined for each bolt diameter, including 12 mm and 16 mm. The aspect ratio of the bolt hole on the connected members is the same as the bolt head to ensure the convenient installation of bolt on one side. Fig. 3 shows two ways of installation in regards to the directions of the long axes of ellipse holes on the cover plates and core plate (i.e. parallel in Fig. 3a and orthotropic in Fig. 3b). Specific dimensions and parameters of the FE models are shown in Fig. 4 and Table 1. Models are named ‘c-d-r’ where ‘c’ represents installation configuration (being P for parallel shown in Fig. 3a, O for orthotropic shown in Fig. 3b), ‘d’ refers to bolt diameter in mm (i.e. 12 or 16 mm) and ‘r’ corresponds to the aspect ratio (L/S) of bolt head ranging from 1.5 to 1.8. 2.2. Numerical investigation Fig. 4a illustrates a typical FE model used in the present study. Each model consists of a one-sided bolt connecting two cover plates and one core plate with the dimensions according to different bolt diameters as shown in Table 1. Half of the double lapped model (two cover plates) was used in the FE model considering the specimen symmetry, where the boundary condition for the symmetric surface of cover plates is set to be fixed according to the symmetry axis. Clearance were considered as

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1 mm for M12 and 2 mm for M16 bolts according to Eurocode 3 part 1e8 [40] (i.e. the short axis of bolt hole on connected plates is 13 mm for M12 joints and 18 mm for M16 joints). The FE model was established and analysed using ABAQUS. For each FE model, threedimensional hexahedral solid element (C3D8) was used for the modelling of the cover plates, core plate and one-sided bolt [41]. The steel plates and bolts were modelled with bilinear stress-strain constitutive relationship. Nominal material properties for G250 steel and G8.8 bolt were used here for determination of aspect ratio (L/S). The material properties for steel plates were defined with elastic modulus (E) of 205 GPa, initial yield stress (s0) of 280 MPa, and Poisson's ratio of 0.3. The yielding of steel is governed by von Mises yield criterion [42]. For the one-sided bolts, elastic modulus (E) was taken as 206 GPa, initial yield stress (s0) 640 MPa (for grade 8.8 bolt) and Poisson's ratio 0.3. The hardening ratio of steel to define the tangent modulus was 1% [43]. Pretension load was simulated by an application of preload force on two parallel surfaces at the middle of bolt shank [44]. In this FE modelling, monotonic displacement was applied to the outer surface of the core plate to induce the shear behaviour of double lapped joint. Contacts in this study mainly include three types: i) the contact between the adjacent surfaces of cover plate and core plate; ii) the contact between bolt head (nut) and cover plate; and iii) the contact between bolt shank and surface of hole walls of cover plates and core plate. For these three types of contacts, behaviours in both tangential and normal directions were considered. For the tangential behaviour, the friction coefficient is isotropic with a value of 0.44 for plate to plate contact and 0.0005 for plate to bolt contact, in accordance to the values reported in references with similar steel surface [45,46]. Hard contact (i.e. penetration between surfaces were limited at the constraint locations and tensile stress was not allowed to be transferred across the interface) was considered for the contact behaviour in the normal direction. 2.3. Comparisons and discussion Two bolt diameters (12 mm and 16 mm) were investigated in this study. For each diameter, assessment of four different aspect ratios of the bolt head and two installation configurations were carried out. The optimal bolt head geometry and installation configuration can then be determined over set of results including yielding capacity, slip distance and stress concentration. In terms of yielding capacity, sixteen cases were analysed including eight joints with M12 bolts and eight joints with M16 bolts. The pretension load for M12 and M16 bolt were 15 and 30 kN respectively. Monotonic displacement was applied to the core plate and the resulting load-deformation curves are shown in Fig. 5. The joints were regarded as reaching their yielding capacity when nonlinearity began to develop during the bearing stage of the loaddisplacement curves as marked in Fig. 5. Fig. 6 shows the stress distribution of connected members of the P-12-1.7 model (i.e. with

Fig. 3. a) Parallel or b) orthotropic installation configuration.

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Table 1 Model dimensions. Bolt type

Model

L/S

Installation configuration

L

S

wp

Core plate lp =tp

lc =tc

M12

P-12-1.5 P-12-1.6 P-12-1.7 P-12-1.8 O-12-1.5 O-12-1.6 O-12-1.7 O-12-1.8 P-16-1.5 P-16-1.6 P-16-1.7 P-16-1.7 O-16-1.5 O-16-1.6 O-16-1.7 O-16-1.8

1.5 1.6 1.7 1.8 1.5 1.6 1.7 1.8 1.5 1.6 1.7 1.8 1.5 1.6 1.7 1.8

Parallel

18 19.2 20.4 21.6 18 19.2 20.4 21.6 24 25.6 27.2 28.8 24 25.6 27.2 28.8

12

72

200/8

96/4

7.5

12

72

200/8

96/4

7.5

16

96

200/12

96/6

10

16

96

200/12

96/6

10

M12

M16

M16

Orthotropic

Parallel

Orthotropic

Cover plate

Head thickness

Fig. 4. Geometric parameters of: a) connected components b) one-sided bolt.

M12 one-sided bolt of 1.7 aspect ratio, Fig. 6a for cover plate, Fig. 6b for core plate and Fig. 6c for one-sided bolt), under the corresponding load (at the marked point of long-dashed curve in Fig. 5a) where nonlinearity initiated. Large area in both cover plate and core plate yielded as shown in red, while the maximum stress on the bolt was still less than the yield stress. Thus, plate bearing failure was found for the joint and in this way the joint yielding capacity can be defined as the onset of the nonlinearity of loaddisplacement curve. Based on the load-deformation curves, the yielding capacity for the four aspect ratios of bolt head in each scenario with certain bolt diameter and installation configuration can be compared to investigate the effect of bolt head aspect ratios. It can be seen that from four curves in each figure in Fig. 5, the yielding capacity for the joints with different bolt head aspect ratios was similar. For example, the yielding capacity was about 80 kN for such M12 bolts with parallel configuration as shown in Fig. 5a (based on the marked points of each curve); the yielding capacity for M12 bolts (with different bolt head aspect ratio) with orthotropic configuration was about the same (80 kN) as depicted in Fig. 5b. Therefore aspect ratio of bolt head did not introduce difference to the joint yielding capacity for a certain diameter of bolts. Through the comparison between Fig. 5a and b or Fig. 5c and d, the installation configuration did not affect the joint yielding capacity neither, i.e. the yielding capacities for P-12-1.5 and O-12-1.5, or P-16-1.5 and O16-1.5, were close to each other. Joints with different bolt diameters obviously presented different yielding capacity, for example 80 kN for M12 from Fig. 5a and b and 165 kN for M16 from Fig. 5c and d. Slip distance is another important consideration for design of

bolted connections. Fig. 7a shows the initial status of the double lapped joint with proposed one-sided bolt before slip. The bolt shank is not in contact with connected plates due to clearance. The slip process can be illustrated in two stages when core plate is subjected to applied shear load (F), i) core plate may slip first until it is in contact with bolt shank (highlighted in blue in Fig. 7b) and then ii) core plate and bolt slip together until the other side of bolt shank contacts with the cover plates (highlighted in red in Fig. 7c). The total distance that core plate travels in these two stages gives the slip distance D as marked in Fig. 7c. The theoretical slip distance is governed by the geometric properties of bolt shank and elliptical hole on connected plates. Fig. 8a shows the relative position between bolt shank and plate hole of the initial stage for parallel installation (top view). By deploying Cartesian coordinate system with origin located at the centre of circular bolt shank (also centre of elliptical hole), a point (xc, yc) at the perimeter of circular bolt shank and a point (xe, ye) at the elliptical plate hole exhibits relationship given in Eqs. (1) and (2) respectively.

x2c þ y2c ¼ r 2 /ð  r < xc < r;  r < yc < rÞ

(1)

where r represents the radius of bolt shank.

x2e y2e þ ¼ 1/ð  a < xe < a;  b < ye < bÞ a2 b2

(2)

where a represents half the length of the major axis and b represents half the length of the minor axis of the ellipse. With a shear load applied to the core plate along the positive x-

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Fig. 5. Load-deformation curves of: M12 a) parallel b) orthotropic configuration M16 c) parallel d) orthotropic configuration.

Fig. 6. P-12-1.7 model von Mises stress contour plots of: a) cover plate b) core plate c) M12 one-sided bolt with 1.7 aspect ratio under joint yielding capacity.

Fig. 7. Schematic connection slip process (not to scale): a) initial stage before slip b) core plate and bolt shank in contact c) core plate and cover plate in contact with bolt shank.

axis, slip of core plate and bolt along positive x-axis occurs. If the core plate slips to a distance of min |xe - xc| in the first stage (see Fig. 8b), the core plate firstly contacts with the bolt shank, completing the slip stage one. Therefore, the slip distance in the first stage can be obtained by combining Eqs. (1) and (2) and solving for min |xe - xc|. For parallel installation of the core plate and cover

plate holes (i.e. Fig. 3a), the slip distance of the core plate together with the bolt shank in the second stage is identical to that received in the first stage, resulting the total slip distance as min 2|xe - xc| (in this case 6.87 mm for P-12-1.7). The theoretical slip distance for orthotropic installation can be calculated using such geometric relationship; and different aspect ratios of bolt head may lead to

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Fig. 8. Relative position between bolt shank and plate hole (parallel installation) a) initial stage b) first stage slip.

different slip distances. The results of theoretical slip distance for each numerical model considering different aspect ratios and bolt diameter are shown in Fig. 9. After the applied shear load overcomes the static friction resistance induced by the applied bolt torque, slip between the core plate and cover plates starts until bearing of the bolt shank and connected plates. The slip process can be identified in the loaddeformation curves where displacement largely increased without load increase. The theoretical values as determined by geometric properties are consistent with those interpreted from FE results in the load-deformation curves from Fig. 5. Furthermore, as shown in Fig. 9, there was an approximately linear relationship between the aspect ratio of bolt head and slip distance; and bolt head with a smaller aspect ratio gave less slip distance. It also can be identified from Fig. 9 that slip distance for orthotropic configuration models reduced by more than 40% in comparison to the corresponding parallel configuration models. This conclusion did not vary with bolt diameters as evidenced in Fig. 9a and b. Therefore, the orthotropic installation configuration may be more preferable than the parallel one considering minimisation of potential slip. Stress concentration on cover plate due to bolt pretension is of further interest in the evaluation of optimal bolt head geometry. As shown in Fig. 10, the von Mises stress on the cover plates for model P-12-1.7 after 15 kN preloading on the bolt may indicate the stress concentration of the cover plate at the bolt head side. Due to the reduced contact area between the bolt head and cover plate, the stress concentration among the contact area may lead to early yielding of cover plate during bolt preloading process. Thus, to investigate the influence of contact area under different bolt pretension level, four different pretension loads (from 15 kN to 30 kN for M12 bolts and 30e45 kN for M16 bolts) were applied to each model and the maximum von Mises stress on the cover plate was recorded. With the increase of the contact area when aspect ratio

increased, the maximum stress on the connected member reduced for a given bolt pretension level as evidenced in Fig. 11. It also illustrates the decreasing trend of the maximum stress when the pretension level reduced for a given model of the double lapped joint. In comparison of four bolt head aspect ratios with M12 parallel configuration in Fig. 11a, the decrease in stress was more apparent from the aspect ratio of bolt head improved from 1.5 to 1.7 and this change became much less when bolt head aspect ratio increased from 1.7 to 1.8. This feature can be identified as well from the models with different bolt diameters and installation configurations as seen in Fig. 11b, c and d. Thus, the alleviation of early yielding induced by the increase of contact area is less effective after bolt head with aspect ratio over 1.7. In addition, a comparison of the models with the same bolt head configuration under the same pretension level (i.e. through the comparison between Figs. 11a and b or Fig. 11c and d) indicates that the maximum stress on cover plates with orthotropic installation configuration was around 20% larger than those with parallel installation configuration. 2.4. Determination of optimal bolt geometry The determination of an optimal bolt head geometry has to consider the outcomes of yielding capacity, slip distance and stress concentration. As shown in Fig. 12, each figure represents one scenario with four different bolt head aspect ratios but a constant bolt diameter (12 or 16 mm) and installation configuration (parallel or orthotropic). In each figure, three axes represent three performance criteria in terms of yielding capacity (C1), slip distance (C2), and stress concentration (C3), with a rating from 0 to 1 where 1 represents the most satisfactory performance (for example the 0.54 for C2, or the 0.92 for C3 for M12 bolt head with 1.7 aspect ratio according to previous evaluations). Lines in different colours stands

Fig. 9. Theoretical slip distance for: a) M12 b) M16 bolt with different bolt head aspect ratios and installation configurations.

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Fig. 10. von Mises stress plots of cover plates for model P-12-1.7 on a) bolt head side b) washer side (with 15 kN bolt preloading).

Fig. 11. Maximum stress plots of four bolt head configurations for: M12 a) parallel b) orthotropic installation configuration; M16 c) parallel d) orthotropic installation configuration.

for different bolt head aspect ratios (i.e. blue stands for bolt head aspect ratio of 1.5). For both M12 and M16 bolt and both parallel and orthotropic configurations, by increasing the aspect ratio from 1.7 to 1.8, the maximum stress on cover plate (C3) only decreased by 8% while the maximum slip distance (C2) increased over 20% and yielding capacity almost remained the same. Therefore, the ellipse bolt head with aspect ratio 1.7 in this study gives the most favourable result considering relatively small clearance and the most effective alleviation for early yield of the cover plates due to bolt preloading. In order to validate the mechanical performance such as yielding capacity and slip distance of one-sided bolted joints with the optimal configuration, an experimental program was

developed and completed for double lapped joints with the proposed one-sided M12 bolt as introduced below. 3. Experimental investigation 3.1. Specimens and materials The M12 one-sided bolt specimens were experimentally examined into two groups considering installation configuration. For both scenarios with parallel and orthotropic configurations, monotonic loading was applied along the long axis of ellipse hole on cover plate (see Fig. 3). A group of ordinary M12 bolt specimens was tested for comparison in addition. The double lapped joint

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Fig. 12. Comparison plots of four bolt head configurations for: M12 a) parallel b) orthotropic installation configuration; M16 c) parallel d) orthotropic installation configuration.

specimens are named ‘DBJ-x’ for this experimental study, where x becomes P for parallel, or O for orthotropic, or N for ordinary bolt configuration. Each group consists of three repeated specimens for data consistency. The plates of the double lapped joint specimens were made of grade 250 steel, where the cover plates are 300  100  4 mm and core plates are 350  100  8 mm as shown in Fig. 13a. Specimen DBJ-N with ordinary bolts had the same dimensions of connected members as the other one-sided specimens. The ellipse holes on steel plates can be produced using laser cutting or hydraulic punching. M12 Grade 8.8 high-tensile ordinary or one-sided bolts were used, and the M12 one-sided bolt were manufactured with the dimensions where the short and long axis of bolt head equals to 12 mm and 20.4 mm respectively (corresponding to the aspect ratio of 1.7), and the shank length equals to the total thickness of connected plates. Clearance 1 mm was prepared for both M12 specimens according to Eurocode 3 [40]. Torque applied on both types of bolts was 36 Nm, resulting in a 15 kN pretension load for these bearing type joints. Coupon tests were conducted following ASTM A370-16 [47] to obtain material properties such as Young's modulus (E), yield strength (fy) and Poisson's ratio (n) of steel plates and the material properties (E, fy and n) of the bolts were provided by the manufacturer as summarised in Table 2. Such material properties and dimensions were also used for the FE modelling in Section 2.2 for following comparison between experimental and FE results. 3.2. Experimental setup and instrumentation Fig. 13a illustrates the positions of strain gauges and string

potentiometer, where three strain gauges (SG2-4) were attached around the bolt hole to detect occurrence of yielding of this region associated with the most remarkable stress concentration during experiments. The other two strain gauges (SG1 and SG5) were positioned further away from the bolt hole to record strain variation. The string potentiometer was fixed onto the specimen within a 200 mm distance between the core and cover plates for relative deformation measurement. Fig. 13b shows a typical experimental setup within the loading machine (Instron600DX). Load was applied to the specimens using displacement control with 0.5 mm/ min. 4. Results and discussion 4.1. Comparison of FE and experimental results The specimens were loaded in shear under displacement control mode until obvious yielding of steel. The yielding capacity was characterised by the nonlinearity in load-displacement curves during the bearing stage, implying the full yielding of the steel plates around the bolt holes. The typical deformed configurations at the bolt holes of cover plates are shown in Fig. 14a for DBJ-P, b for DBJ-O and c for DBJ-N, where permanent deformation of the steel plate around the ellipse (one-sided bolt specimens) or circular (ordinary bolt specimens) hole was noted in each cover plate due to the contact with the bolt shank and the resulted bearing behaviour. DBJ-P (Fig. 14a) and DBJ-O (Fig. 14b) exhibited similar deformation there due to their identical cover plate orientation. The experimental load-deformation curves of all the specimens

C. Wan et al. / Journal of Constructional Steel Research 165 (2020) 105815

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Fig. 13. a) Dimensions of DBJs b) Experimental set up.

Table 2 Material properties for assemblies. Material

E (GPa)

n

fy (MPa)

fu (MPa)

Cover plate Core plate G8.8 bolt

200.98 201.62 206

0.282 0.286 0.3

314.67 311.93 740

458.53 455.26 840

are shown in Fig. 15 together with the corresponding FE results. Such responses can be seen in three stages in general. At first, the applied load was less than the static friction force between the connected members due to the applied pretension force through the bolt torque. There was a linear increase of load before the initial slip point A as marked in each figure. When the applied load overcame the static friction, the contact surfaces started to slip until

Fig. 14. Deformation on cover plate of a) DBJ-P b) DBJ-O c) DBJ-N.

Fig. 15. Load-deformation curves for: a) DBJ-P b) DBJ-O c) DBJ-N.

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C. Wan et al. / Journal of Constructional Steel Research 165 (2020) 105815

bearing occurs (point B in Fig. 15). At this stage, the applied load maintained approximately at the level of 14 kN while the relative displacement between the core and cover plate increased significantly (from point A to B as seen in Fig. 15 due to such slip). The maximum slip distance is the clearance of bolt hole as governed by the geometric properties (see Section 2.3). After both two sides of the bolt shank contacted with the connected cover plates and core plate, the bolt shank started to squeeze the connected plates with the increase of applied load, i.e. the bearing stage starts (after point B). The specimens were regarded as reaching their joint yielding capacity where nonlinearity began to develop in the curves as marked in Fig. 15 (solid rhombus for FE results and triangle for experiment results). The FE stress distribution on the bolts under the joint yielding capacity of specimens are shown in Fig. 16. For all specimens, the bolts did not yield under the joint yielding capacity as the maximum stresses shown were less than the bolt yield stress 740 MPa. As depicted in Fig. 17, a large area around the bolt hole on the connected plates yielded (shown in red) for the specimens under joint yield capacity. Therefore, the joint yielding capacity corresponded to substantial yielding of the connected members. On comparison of the load-displacement curves between experimental and FE results in Fig. 15, three main aspects can be identified including the static friction resistance induced by bolt torque (point A), the slip distance (stage AB) and the joint yielding capacity (marked points). As can be seen from the FE results in Fig. 15, the friction resistance for specimens with one-sided bolt (DBJ-P and DBJ-O) and ordinary bolt (DBJ-N) were about 14 kN. Even though the contact area between bolt head and cover plates were different for one-sided and ordinary bolted joints, the consistent friction resistance indicates that friction resistance between connected plates only related to number of contact surfaces (as 2 here for double lapped joint), the friction parameter (m ¼ 0.44) for plate contact and applied bolt pretension (P ¼ 15 kN). This is consistent with Eq. (3) as concluded in Ref. [48]. The calculated friction resistance is 13.2 kN. Experimental, FE and theoretical results show satisfactory agreement with only slight difference.

ordinary bolted specimens (DBJ-N) with designated 1 mm clearance as shown in Fig. 15d, the results were more affected by the installation, manufacturing and measuring accuracies, where the overall discrepancy is about 68%. From the load-deformation curves of DBJ-P specimens with onesided bolts and parallel configuration as shown in Fig. 15a, the FE modelling predicted an earlier appearance of nonlinearity at 90.2 kN (joint yielding capacity), than the experimental results of 99.6 kN. The joint yielding capacity (both experimental and FE results) for other specimens (DBJ-O and DBJ-N) were interpreted from Fig. 15 and summarised in Table 3. Similar as DBJ-P, the predictions of joint yielding capacity for DBJ-O and DBJ-N from FE analysis were lower than those obtained from experimental results. A likely reason is the approximations in the modelling of material properties. The maximum difference of joint yielding capacity between experimental (Fy,E) and FE results (Fy,FE) are within 16.5%, implying the nonlinear bearing behaviour can be successfully captured by FE modelling. The load-strain responses were further plotted in Fig. 18 for certain locations of the cover plate (see Fig. 13a). Identical loadstrain responses were identified for SG2 and SG4 due to symmetry location, thus only SG2 was plotted. From the load-strain curves of DBL-P shown in Fig. 18a, SG2 exhibited nonlinearity earlier than SG3, suggesting the yielding initiation on cover plates, at 63 kN (Fi,FE) according to the FE results, or 61 kN (Fi,E) according to experimental results. Depicted in Fig. 18b, DBJ-O exhibits similar load-strain response as DBJ-P, with yielding initiation at the location of SG2 predicted by FE modelling at 60 kN (Fi,FE), or experimental results of 63 kN (Fi,E). DBJ-N experienced yielding initiation of cover plate at around 40 kN from both experimental and FE results at the location of SG3 as shown in Fig. 18c. No yielding was observed from the locations of SG1 and 5. Listed in Table 3, the discrepancy between FE and experimental results of yielding initiation are within 5%.

Fn ¼ nmp

As depicted in Fig. 15a, the slip distance (SE) for specimens DBJ-P (parallel configuration) is 6.2 mm from experimental results. Specimens DBJ-O was with identical long axis loading as specimens DBJ-P but orthotropic configuration, the slip distance (SE) can be reduced significantly from 6.2 mm to 2.23 mm as shown in Fig. 15b. The initial contact for both specimens DBJ-P and DBJ-O included two points as a result of the ellipse bolt hole configuration (see points highlighted in red in Fig. 19a). As indicated in Fig. 18, the initial yielding of cover plate started from the location of SG2 (and therefore also SG4 because of symmetry) for specimens DBJ-P and DBJ-O where the SG2 curves deviated from initial linearity earlier than SG3. This better explains the identical initial contact area of the two scenarios with one-sided bolts, therefore similar initial bearing stiffness of joints on bearing stage was observed from the

(3)

where, n is the number of contact surface between connected plates, m is friction parameter between plates and P is the pretension load applied to the bolt. The maximum slip distance depends on the geometric relationship between the bolt holes of connected plates and bolt shank, theoretical values as determined in Section 2.3 can be used for comparison with experimental results as shown in Table 3 for the specimens in three scenarios. The slip distance of one-sided specimens from experiments is less than those from theoretical values, with the maximum percentage difference less than 40%. Such differences were related to the unavoidable manufacturing inaccuracies of the connected plates and installation precision. For the

4.2. Effects of installation configurations

Fig. 16. Bolt von Mises stress plots for a) DBJ-P b) DBJ-O c) DBJ-N under joint yielding capacity.

C. Wan et al. / Journal of Constructional Steel Research 165 (2020) 105815

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Fig. 17. von Mises stress distribution at plate bearing location for a) DBJ-P b) DBJ-O c) DBJ-N under joint yielding capacity.

Table 3 Specimens and their yielding capacities and slip distances. Specimen groupa

Yielding initiation loadb (kN) Fi,E

Fi,FE

Fi,E/Fi,FE

Fy,E

Fy,FE

Fy,E/Fy,FE

SE

Sth

DBJ-P DBJ-O DBJ-N

61 63 42

63 60 44

0.968 1.05 0.954

99.6 98.3 101.2

90.2 86.7 86.8

1.104 1.133 1.165

6.2 2.23 0.32

6.87 3.64 1

a b c d

Yielding capacityc (kN)

Slip distanced (mm)

Specimens named following ‘DBJ - installation arrangement. Point corresponding to first point on connected plate reached yield strain (0.1517%). Intersection point of the load-deformation curve's elastic and post-yield tangent lines. Dominated by geometry property, compared between experimental and theoretical values.

load-deformation curves in Fig. 15. The initial yielding of cover plate started at loads 61 kN (Fi,E) for specimens DBJ-P and 63 kN (Fi,E) for specimens DBJ-O with orthotropic installation (see Table 3). The loads of initial yielding for specimens DBJ-P and DBJ-O are similar. Again, this is because the same initial contact behaviour of specimens DBJ-P and DBJ-O resulting from identical orientation of ellipse bolt hole on cover plates. The failure mode of specimens in each scenario was identified as excessive yielding of the cover plates around the bolt hole. As

indicated in Table 3, the yielding capacity for specimens DBJ-P (with parallel configuration) is almost the same (about 99 kN) as that of specimens DBJ-O (with orthotropic configuration), due to the same orientation of cover plate and identical plate yielding failure mode as shown in Fig. 14a and b. This is because the yielding capacity relates to the yielded area around the bolt hole. As shown in Fig. 17, the yielded area shown in red for specimens DBJ-P (Fig. 17a) appeared to be similar to those of DBJ-O specimens (see Fig. 17b). Due to the identical orientation of ellipse hole on cover

Fig. 18. Load-strain responses from the strain gauges on the cover plate: a) DBJ-P b) DBJ-O c) DBJ-N.

Fig. 19. The relative position between bolt shank and plates a) Before slip b) After bearing.

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C. Wan et al. / Journal of Constructional Steel Research 165 (2020) 105815

plate, the deformed area (enclosed by grey shade in Fig. 19b) for DBJ-P (with parallel installation) is similar to those of specimens DBJ-O (with orthotropic installation), resulting their similar yielding capacities. 4.3. Comparison with ordinary bolts It has been shown in Fig. 15 that the slip distance (SE) with a value of 0.32 mm for specimen DBJ-N (with ordinary bolts) is less than other two scenarios with one-sided bolts (6.2 mm for DBJ-P and 2.23 mm for DBJ-O). The yielding capacity for specimens DBJN with ordinary bolts exhibited only minor difference to specimens DBJ-P and DBJ-O (see Table 3). This is because of the similar yielded area on cover plates as can be recognised in Fig. 17. In comparison to the specimens with one-sided bolts, DBJ-N with ordinary bolts presented about 24% higher initial bearing stiffness according to the load-displacement curves in Fig. 15. This is induced by the different initial contact area and radius along the contact surface depicted in Fig. 19a. From the load-strain responses shown in Fig. 18, the yielding initiation load (Fi,E) for specimen DBJN (with ordinary bolt) was 42 kN and less than those of the onesided specimens (around 60 kN) started at the location of SG3. This is because initial single point to point contact (at the location of SG3) was introduced for specimens DBJ-N while two-point initial contact (at the locations of SG2 and SG4) was observed for specimens DBJ-P and DBJ-O as indicated in Fig. 19a. 5. Conclusions An innovative one-sided bolt with ellipse bolt head was proposed and examined in this paper. Numerical analysis was carried out first to determine the optimal head geometry considering joint yielding capacity, slip distance and stress concentration. Experimental and FE modelling were further conducted to investigate the shear performance of double lapped joints with the proposed onesided bolt and to understand the effect of different installation configurations. A group of ordinary bolted joints was studied for comparison in addition. Based on this work, the following conclusions can be drawn: 1. The aspect ratio of ellipse bolt head investigated in this study, as well as parallel or orthotropic installation configuration, did not affect much the yielding capacity of the double lapped joints with the proposed one-sided bolts. However, the slip distance increased linearly with the increase of ellipse bolt head aspect ratio as this is dominated by the geometric relationship between bolt hole on connected members and bolt shank. With the increase in aspect ratio of the ellipse bolt head, the stress concentration due to bolt preloading on the bolt head side was alleviated. The ellipse bolt head with aspect ratio 1.7 gives favourable results in this study considering relatively small slip distance and the most effective alleviation for early yielding of the cover plate. 2. The static friction resistance is not affected by the contact area between bolt head and connected plates but related to the number of contact surface, friction parameter between connected plates and applied bolt pretension only. The slip distance was governed by the geometric characteristics of the bolted connection such as bolt hole aspect ratio and such experimental results may also be affected by the manufacture and installation accuracy. 3. When the load-deformation curve began to exhibit nonlinearity during bearing stage, the joint reached its yielding capacity as the area around bolt hole experienced excessive yielding through the observation of local strain responses. Specimens DBJ-P (parallel configuration) and DBJ-O (orthotropic

configuration) exhibited similar yielding capacity (discrepancy within 3%) in comparison to specimens DBJ-N with ordinary bolt, evidenced by the load-deformation behaviour from both FE and experimental results. The FE predictions of joint yielding capacity were within 16.5% difference in comparison to the experimental ones. 4. The load-strain curves from experiments and FE modelling compared well, again validating the FE modelling in prediction of the yielding initiation on the bolted joint specimens. The yielding initiation was related to initial contact between bolt shank and bolt hole. Specimens DBJ-P and DBJ-O with identical cover plate orientation, showed similar initial contact behaviour (two-point contact), resulting in 33% higher load for yielding initiation than that from specimens DBJ-N with single point to point initial contact. 5. Specimens DBJ-O (with orthotropic configuration) exhibited comparable load for yielding initiation and yielding capacity in comparison to specimens DBJ-P (with parallel configuration). The slip distance of specimens DBJ-O was 64% less than those of specimens DBJ-P, which is more preferable considering the minimisation of potential slip. Accordingly, the orthotropic installation configuration may be a preferred approach for implementation of such one-sided bolts in shear. The orthotropic installation can be done by first inserting the bolt head into the ellipse hole and then aligning with the orthotropic ellipse hole on the other member. After the bolt head passes through both the members, the bolt is rotated to be orthotropic to the next ellipse hole after passing through and then tightened up. The slip distance of specimens DBJ-O with orthotropic configuration was 2.23 mm and yield capacity was 98.3 kN, in comparison to 0.32 mm and 101.2 kN for the corresponding joint specimens DBJ-N with ordinary bolts. Acknowledgements The authors acknowledge the support from the Australian Research Council through the Discovery Project (DP180102208). The authors also gratefully acknowledge the staff in the Civil Engineering Laboratory of Monash University Australia for their assistance in the experiments. References [1] X.-L. Zhao, G.J. Hancock, Square and rectangular hollow sections subject to combined actions, J. Struct. Eng. 118 (1992) 648e667. [2] X.-L. Zhao, G.J. Hancock, Square and rectangular hollow sections under transverse end-bearing force, J. Struct. Eng. 121 (1995) 1323e1329. [3] Y. Liu, B. Young, Buckling of stainless steel square hollow section compression members, J. Constr. Steel Res. 59 (2003) 165e177. [4] F.J. Luo, C. Ding, A. Styles, Y. Bai, End plateestiffener connection for SHS column and RHS beam in steel-framed building modules, International Journal of Steel Structures 19 (2019) 1353e1365. [5] J. Wardenier, J. Packer, X. Zhao, G. Van der Vegte, Hollow Sections in Structural Applications, Bouwen met Staal Rotterdam, The Netherlands, 2002. [6] D. Ridley-Ellis, G. Davies, J. Owen, Fundamental Behaviour of Rectangular Hollow Sections with Web Openings, Eighth International Symposium on Tubular Structures, Singapore, 1998. [7] Y. Ding, L. Zhu, K. Zhang, Y. Bai, H. Sun, CHS X-joints strengthened by external stiffeners under brace axial tension, Eng. Struct. 171 (2018) 445e452. [8] C.-C. Chen, C.-C. Lin, C.-L. Tsai, Evaluation of reinforced connections between steel beams and box columns, Eng. Struct. 26 (2004) 1889e1904. [9] C. Ding, X. Pan, Y. Bai, G. Shi, Prefabricated connection for steel beam and concrete-filled steel tube column, J. Constr. Steel Res. 162 (2019) 105751. [10] B. Yang, K.H. Tan, Numerical analyses of steel beamecolumn joints subjected to catenary action, J. Constr. Steel Res. 70 (2012) 1e11. [11] W. Tizani, A. Al-Mughairi, J.S. Owen, T. Pitrakkos, Rotational stiffness of a blind-bolted connection to concrete-filled tubes using modified Hollo-bolt, J. Constr. Steel Res. 80 (2013) 317e331. [12] L. Deng, W. Yan, L. Nie, A simple corrosion fatigue design method for bridges considering the coupled corrosion-overloading effect, Eng. Struct. 178 (2019) 309e317. [13] A.J. Hoogenboom, Flow Drill for the Provision of Holes in Sheet Material, Google Patents, 1984.

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