Structures and Buildings Volume 168 Issue SB10 Pull-out behaviour of blind bolts from concrete-filled tubes Oktavianus, Yao, Goldsworthy and Gad

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1 Volume 168 Issue SB1 Proceedings of the Institution of Civil Engineers Structures and Buildings 168 October 215 Issue SB1 Pages Paper 1498 Received 3/1/214 Accepted 26/2/215 Published online 2/7/215 Keywords: anchors & anchorages/composite structures/ engineering mechanics ICE Publishing: All rights reserved Pull-out behaviour of blind bolts from &1 Yusak Oktavianus Msc (Eng) PhD candidate, Department of Infrastructure Engineering, University of Melbourne, Melbourne, Australia &2 Huang Yao PhD Managing Director, Australia Xing One Pty Ltd, Lidcombe, Australia &3 Helen M. Goldsworthy PhD Associate Professor, Department of Infrastructure Engineering, University of Melbourne, Melbourne, Australia &4 Emad F. Gad PhD Professor, Chair of Civil and Construction Engineering, Faculty of Science, Engineering and Technology, Swinburne University of Technology, Hawthorn, Australia The use of concrete-filled steel circular hollow sections as columns is becoming more popular because of their superior capacity, good ductility and large energy absorption capacity. The utilisation of blind bolts, which can be installed from the outside of the column, is possible and has been the subject of considerable research. Results from pull-out testing of blind bolts and the headed anchor blind bolt, which is a modification of the original blind bolt, are reported in this paper. The tests were performed to establish the behaviour of individual bolts. Results from the research will ultimately be used in the design of moment-resisting connections using these bolts. The effect of varying several parameters such as the tube thickness, bolt diameter and embedment depth were investigated both experimentally and by way of numerical models using finite-element analysis. After achieving a good agreement between the experimental and numerical results, further analysis was implemented to determine the relative contributions of the concrete and tube wall to the pull-out resistance. Parametric studies on concrete strength and embedment depth were also performed. 1. Introduction The use of concrete-filled (CF) circular hollow sections (CHSs) as columns is becoming more popular because of their superior capacity, good ductility and large energy absorption capacity under seismic action (Han and Li, 21). The concrete infill assists the CHS in resisting buckling, and the CHS provides confinement for the concrete. Other benefits of the use of CFCHS columns are fast construction, as they do not need additional formwork, and aesthetic appeal. Conventional bolts cannot be used to connect to a CHS because of lack of access to the inside of the section. This problem has led to the development of blind bolts, which can be installed from the outside of the CHS. The commercially available blind bolts include the Huck high-strength blind bolt (Huck International, 199), the Lindapter Hollo-bolt (Lindapter International, 1995), flow drilling (France et al., 1999) and the Ajax Oneside (Ajax Engineered Fasteners, 22). There are different types of blind bolts on the market, and this paper focuses on one type, to demonstrate the mechanical behaviour of a system rather than the performance of a specific blind bolt. The Ajax Oneside blind bolt was used in this research, and is referred to throughout this paper as BB. Research on BB connections to unfilled Square hollow section (SHS) columns was carried out by Lee et al. (21, 211a, 211b). Taking advantage of the concrete infill, the BB can be modified either using a cogged anchor (CABB) (Yao et al., 28) or headed anchor (HABB) (Yao et al., 211). An experimental and numerical investigation of the pull-out behaviour of CABBs in CFCHSs has been reported in Yao et al. (28). Although the CABBs provided good strength and stiffness, HABBs are considered to be more practical than CABBs because HABBs can be easily manufactured in one piece. Furthermore, disorientation of the cogged anchor, which sometimes occurred during the installation of the CABBs, did not occur if HABBs were being used. A similar concept utilising additional embedment depth has also been investigated by Mahmood et al. (214) and Pitrakkos and Tizani (213) on an Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 747

2 Volume 168 Issue SB1 extended Hollo-bolt in a CFSHS and by Agheshlui (214) on HABBs in CFSHSs. Although experimental results from pullout tests of single BBs and HABBs in CFCHSs have been reported in Yao et al. (211), a numerical study has not previously been carried out to investigate the important parameters in this case. This paper focuses on numerical analyses investigating the pull-out behaviour of single BBs and HABBs embedded in CFCHSs with D/t ratios from 32 4 to 54, where D is the outer diameter of the CHS of a bolt and t is the thickness of its CHS. Investigating the full range of the tensile behaviour of single BBs and HABBs embedded in CFCHSs is the main aim of this research. (The behaviour of connections under combined actions (i.e. shear and tension) is still under investigation, and is not presented in this paper.) Figure 1 shows the components of the HABB. Several parameters, such as CHS thickness, diameter of the HABB and embedment depth, were varied both in the experimental work and in the finiteelement analyses (FEAs). Of particular interest in this study is the pull-out capacity of the bolts in the different cases under consideration and whether it is sufficient to ensure that the full tensile capacity of the bolt is reached. The secant stiffness at 6% of the nominal ultimate capacity is also considered, because it is a key variable for design. This research has three important objectives, as follows. Firstly, finiteelement models were constructed using ABAQUS, and the results were compared with the experimental findings. The results show that good agreement between the experimental and FEA results has been established for the pull-out force compared with the displacement behaviour. Second, further analysis using FEAs in determining the relative contribution of the bearing of the headed anchor on the concrete (concrete contribution) and the bearing of the washer on the steel tube wall (steel contribution) to the pull-out resistance was implemented. Third, parametric studies on the concrete strength and embedment depth were also performed, leading to a proposal for the minimum required concrete strength and embedment depth for the specimens considered in this research. Bolt head HABB Collapsible washer Solid washer Bolt head Sleeve Nut 2. Experimental work 2.1 Experimental set-up Static tests on individual blind bolts in tension were conducted to obtain the pull-out force in comparison with the outward displacement. The results will be used as the basis to determine the stiffness and the strength of the connection. The details of the experimental set-up were explained in Yao et al. (211). In this paper, some important information is repeated briefly. The set-up of the experiment is illustrated in Figure 2. The thrust blocks shown in Figure 2 were placed on both sides at a distance of 2 mm from the centre of the tube hole (which is equal to two times the embedment depth, to ensure that the formation of the concrete cone would not be affected by the boundary conditions). Two inner linear variable differential transformers (LVDTs) (T2 and T3) and two outer LVDTs (T1 and T4) were installed in the flat plate welded to the nut and mm away from the perimeter of the hole, respectively (Figure 2). The outer LVDTs were installed to measure the total deformation of the blind bolt and the bar over the predefined length outside the CHS. The inner LVDTs were installed to measure the stretching of the bar over the predefined length outside the CHS. Therefore, the outward displacement of the blind bolt was calculated by subtracting the average measurements of inner LVDTs (T2 and T3) from the average measurements of the outer LVDTs (T1 and T4). The outer diameter of the CHS (D) was 324 mm and the thicknesses (t) were 6, 8 and 1 mm. The grade of the CHS was C35L. The density of the concrete was 234 kg/m 3, with a slump of 1 mm. Each cubic metre of concrete consisted of kg of water, kg of cement as per AS 3972 (Standards Australia, 21), kg of Lyndhurst sand, and 351 and 819 kg of 1 and 2 mm sandstone crushed aggregate, respectively. The mean compressive strength, from three CFCHS b T1 a C T2 HABB T3 T4 LVDT Hydraulic jack Clamp Load cell Reaction frame Thrust block Embedment depth t D a = mm b = 2 mm C = flat plate welded to the nut Figure 1. HABB components Figure 2. The experimental set-up 748 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved.

3 Volume 168 Issue SB1 cylinder tests of concrete, was 48 MPa. The BB and HABB were grade 8 8. The BB material properties and other details can be found in an Ajax technical note (Fernando, 25). The details of the specimens used in the experiment are shown in Table 1. The notation used is as follows: T6 denotes a specimen with a tube thickness of 6 mm, D16 denotes a specimen with a blind bolt diameter of 16 mm, N1 indicates a normal blind bolt (i.e. a BB), N2 indicates a HABB with an embedment depth of 1 mm, and E indicates experimental work. The embedment depth in this paper is measured along the blind bolt from the beginning of the first head to the beginning of the second head, as shown in Figure 1. A rounded nut was used as the head, and both the first and second heads have the same thickness. Therefore, the effective embedment depth will be equal to the sum of the embedment depth and the thickness of the washer. The bolts were tightened to a snug tight condition, defined as the tightness achieved with a few impacts of an impact wrench or by the full effort of a person using a standard podger spanner (Standards Australia, 1998). The purpose of snug tight is to bring the connected plies into firm contact (AISC, 21). 2.2 Experimental results The relationship between the pull-out force and the outward displacement was used to evaluate the tensile performance of the BB and HABB. Figure 3 shows the relation between the pull-out force and the outward displacement of the blind bolts. The outward displacement of the blind bolts was obtained by subtracting the average outward displacement of LVDTs T2 and T3 from that of LVDTs T1 and T4, as explained previously. In the case of N1, the failure mechanisms were yielding of the tube wall followed by pull-out of the blind bolt, except for specimen T8_D16_N1_E, which failed by bolt fracture. On the other hand, the failure mechanism for all of the N2 cases (Figure 3(b)) was blind bolt fracture. Figure 3(a) shows that increasing either the CHS thickness or the blind bolt diameter increased both the stiffness and the pull-out capacity. In the case of N1, increasing the thickness of the tube wall is more effective in increasing both the stiffness and strength than increasing the diameter of the blind bolt (Figure 3(a)). The reverse phenomenon was observed in the case of N2 (Figure 3(b)): that is, increasing the diameter of the blind bolt is more effective in increasing both the stiffness and strength than increasing the Specimen Tube size D/t Bolt diameter: mm Embedment depth: mm T6_D16_N1_E CHS T6_D2_N1_E CHS T8_D16_N1_E CHS T8_D2_N1_E CHS T1_D2_N1_E CHS T6_D16_N2_E CHS T6_D2_N2_E CHS T8_D16_N2_E CHS T8_D2_N2_E CHS T1_D2_N2_E CHS Table 1. Details of the specimens in the experiment T6_D16_N1_E T6_D2_N1_E T8_D16_N1_E T8_D2_N1_E T1_D2_N1_E T6_D16_N2_E T6_D2_N2_E T8_D16_N2_E T8_D2_N2_E T1_D2_N2_E (a) (b) Figure 3. Experimental results: (a) case N1; (b) case N2 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 749

4 Volume 168 Issue SB1 thickness of the tube wall. Further explanation of this behaviour is presented in subsequent sections on the results of the FEAs. 3. Finite element analysis 3.1 Finite element modelling ABAQUS/explicit was used to perform the FEAs. As an illustration, the ABAQUS model of specimen T6_D16_N2_F (where F indicates FEA) is shown in Figure 4. Taking advantage of the axes of symmetry, a quarter model was used. A displacement boundary condition was applied to the tip of the extended blind bolt, to represent the action of the hydraulic jack as displacement controlled loading. The following assumptions were made to establish the finite-element model. The bolt and nut were modelled as one solid element. It was assumed that there is no pretension force on the bolt. The collapsible washer was also modelled as a solid washer. The C3D8R element type (three-dimensional (3D) stress, eight-node linear brick, reduced integration) was used for all elements, except a small partition (one-quarter of circle) in the centre of the blind bolt, for which C3D6 (3D stress, six-node linear triangular prism) was chosen to produce a uniformly distributed mesh along the bolt section. This reduced the computational time and increased the accuracy. The size of the mesh was chosen carefully, and a reduction in the size of the mesh does not cause a significant change in the result. General contact with hard contact in normal behaviour and a friction penalty in tangential behaviour were used. The coefficient of friction used between the concrete and the HABB was equal to 3, as suggested by Guezouli and Lachal (212). 3.2 Material properties Nominal mechanical properties were used for the steel tube, blind bolt, nut and washer. The steel tubes were produced in accordance with AS 1163 (Standards Association of Australia, 1991), and have a grade of 35L. On the other hand, the blind bolt, nut and washer were manufactured as per AS (Standards Australia, 2) for high-strength bolts, and have a grade of PC8 8. Multi-linear stress strain curves, as summarised in Table 2, were used to obtain better approximations. The Mander model (Mander et al., 1988) was used to model the stress strain curve of the unconfined concrete. The mean compressive strength consistent with the experimental results of 48 MPa was used to represent the actual compressive strength of the concrete. The calculation of the modulus of elasticity of the concrete was based on a relationship given in the Australian standard for concrete structures (Standards Australia, 29). Concrete damage plasticity was used for modelling the concrete plastic behaviour. The fracture energy, which is approximately equal to 4 N/m for a mean compressive strength of concrete ( f cm ) equal to 2 MPa, and N/m for f cm equal to 4 MPa, was used for modelling the concrete behaviour in tension following the ABAQUS documentation Steel tube Nut Washer Washer Bolt head HABB Concrete infill Bolt head (a) (b) Figure 4. Finite-element model of specimen T6_D16_N2_F: (a) quarter model; (b) zoom-in view 75 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved.

5 Volume 168 Issue SB1 Note Fracture energy: N/m Steel tube Blind bolt, nut and washer Stress: MPa Strain Stress: MPa Strain Initial Yield point Post-yield Post-yield Ultimate Table 2. Material properties for the steel tube and blind bolt, nut and washer Concrete mean compressive strength, f cm : MPa Figure 5. Relationship between the fracture energy and concrete compressive strength (ABAQUS, 211). For the concrete strength of 48 MPa, linear extrapolation was used, and resulted in a fracture energy approximately equal to 15 N/m (Figure 5). The tensile strength of the concrete was assumed to be approximately equal to p 56 ffiffiffiffiffiffiffi f cm, which is similar to that specified in ACI 318M (ACI, 211). 3.3 FEA result The pull-out force and outward displacement results obtained from the FEAs are shown in Figure 6. The behaviour is similar to that observed experimentally. Comparisons between these results and the experimental results are given in the next section. By adding the headed anchor, the full tensile capacity of the blind bolt is achieved. Figure 6(a) shows that, in the case of N1, increasing the thickness of the tube wall is more effective in increasing both the stiffness and strength than increasing the diameter of the blind bolt, as was also observed experimentally. This is because in the range of available tube wall thicknesses, the deformability and strength of the tube wall dominates the behaviour. Figure 6(b) shows that, in the case of N2, increasing the thickness of the tube wall is less effective in increasing both the stiffness and strength than increasing the diameter of the blind bolt. This is because the behaviour of the anchorage is dominating in the cases with sufficient embedment depth. However, one of the FEAs, T1_D2_N1_F, predicts a different mechanism to that observed in the test: blind bolt fracture rather than the pull-out failure observed in the experimental result. This is simply because the nominal capacity of the bolts used in the FEAs is successfully reached prior to pull-out, whereas in the experiment the bolts actually have a higher capacity, and this cannot be reached. 4. Comparison between the experimental and finite-element results Comparisons between the experimental and the finite-element results are plotted in Figure 7 and summarised in Table 3. In a T6_D16_N1_F T6_D2_N1_F T8_D16_N1_F T8_D2_N1_F T1_D2_N1_F T6_D16_N2_F T6_D2_N2_F T8_D16_N2_F T8_D2_N2_F T1_D2_N2_F (a) (b) Figure 6. Finite-element results: (a) case N1; (b) case N2 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 751

6 Volume 168 Issue SB T6_D16_N1_E 6 6 T6_D2_N1_E T6_D16_N1_F 4 T6_D16_N2_E 4 T6_D2_N1_F 2 Collapsible washer fails T6_D16_N2_F_1 1tube Collapsible washer fails T6_D16_N1_F_1 1tube 2 T6_D16_N2_F 2 T6_D2_N1_F_1 1tube (a) (b) (c) T6_D2_N2_E T6_D2_N2_F T8_D16_N1_E T8_D16_N2_E T6_D2_N2_F_1 1tube T8_D16_N1_F T8_D16_N2_F (d) (e) (f) T1_D2_N1_E T8_D2_N1_E 4 T8_D2_N2_E 4 2 T1_D2_N1_F T8_D2_N1_F T8_D2_N2_F T1_D2_N1_F_1 25bb (g) (h) (i) Figure 7. Comparison between the experimental and FEA results: (a) T6_D16_N1; (b) T6_D16_N2; (c) T6_D2_N1; (d) T6_D2_N2; (e) T8_D16_N1; (f) T8_D16_N2; (g) T8_D2_N1; (h) T8_D2_N2; (i) T1_D2_N1; ( j) T1_D2_N2 T1_D2_N2_E 4 T1_D2_N2_F T1_D2_N2_F_1 25bb (j) 752 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved.

7 Volume 168 Issue SB1 Specimen Failure mode (FEA) Ultimate strength, F u :kn Ultimate strength error: % 6F u : kn a Secant stiffness at 6F u, K 6 : kn/mm Secant stiffness error: % T6_D16_N1_F Tube wall yield and pull-out T6_D16_N1_F_1 1tube Tube wall yield and pull-out T6_D16_N2_F Bolt fracture T6_D16_N2_F_1 1tube Bolt fracture T6_D2_N1_F Tube wall yield and pull-out T6_D2_N1_F_1 1tube Tube wall yield and pull-out T6_D2_N2_F Bolt fracture T6_D2_N2_F_1 1tube Bolt fracture T8_D16_N1_F Tube wall yield and bolt fracture T8_D16_N2_F Bolt fracture T8_D2_N1_F Tube wall yield and pull-out T8_D2_N2_F Bolt fracture T1_D2_N1_F Tube wall yield and bolt fracture T1_D2_N1_F_1 25bb Tube wall yield and pull-out T1_D2_N2_F Bolt fracture T1_D2_N2_F_1 25bb Bolt fracture a F u is taken as the ultimate strength of the FEA results using steel nominal mechanical properties. Table 3. Strength, failure mode and secant stiffness beam column connection that uses HABBs, the blind bolts would not be allowed to exceed approximately 6% of their nominal ultimate capacity. The reason for this is that, as the load is increased above this level, micro-cracks continuously develop around the embedded head, and this results in stiffness degradation. Hence, it is useful to assess the secant stiffness at this force level. The secant stiffnesses (at 6% of the nominal ultimate capacity) of the specimens from the experiment and FEAs are compared in Table 3. Table 3 shows the maximum error in the secant stiffness is about 32% for specimen T6_D16_N1_F. This large error could be due to the use of the nominal steel yield strength in the FEA. The material properties of the steel tube have been altered to illustrate the sensitivity of the results to these parameters. If, for example, the yield and ultimate strengths of the 6 mm thick steel tube were increased by 1% (designated as 1 1tube in Figure 7(a) 7(d)), the error in the secant stiffness would generally be reduced, sometimes by a large margin, but in other cases with little change. However, the error in the ultimate strength of the specimens is also sensitive to the steel tube material properties, and, in some of the cases considered here, the 1% increase in the yield and ultimate tensile strength of the steel tube material has led to an increase in the error in the ultimate strength of the specimen. For example, in the case of T6_D16_N1_F there is a dramatic decrease in the error in the secant stiffness (from 32% for T6_D16_N1_F to only 6% for T6_D16_N1_F_1 1tube (Figure 7(a) and Table 3)), but a slight increase in the error in estimating the ultimate strength of the specimen (from 4% to 8%). Other material properties that could be substantially different to the assumed nominal values are the yield and ultimate tensile strength of the blind bolt, nut and washer. This is because the nominal ultimate tensile strength, for example, is a characteristic value that is exceeded by at least 95% of the actual bolts. If, for example, the yield and ultimate strengths of the blind bolt, nut and washer for the cases of D2 bolts in 1 mm tubes were increased by 25% (designated as 1 25bb in Figure 7(i) and 7( j)), the error in strength would generally decrease. However, the error in the secant stiffness is also sensitive to the material properties of the blind bolt, and in both of the cases considered here there is an increase in this error due to the assumed 25% increase in the yield and ultimate tensile strength of the blind bolt, nut and washer. In the HABB case of T1_D2_N2_F, there is a large decrease in the error in the ultimate strength (from 2% for T1_D2_N2_F to 4% for T1_D2_N2_F_1 25bb (Figure 7( j) and Table 3)) and a small increase in the error in the secant stiffness (from 1% for T1_D2_N2_F to 1 5% for T1_D2_N2_F_1 25bb (Figure 7(j) and Table 3)). From the above discussion, it is clear that small realistic increases in the yield and ultimate strengths of the steel tube Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 753

8 Volume 168 Issue SB1 could reduce the error in secant stiffness; while similar increases in the yield and ultimate tensile strengths of the blind bolt, nut and washer could reduce the error in the ultimate strength. One anomaly appears to be that the FEAs for specimens T6_D16_N1_F and T6_D2_N1_F overestimate the ultimate capacity by 4% and 13%, respectively. This occurs because the failure mechanism in the washer is not captured in the FEAs. In the experiment the washer failed, and the load could not increase further (Figure 7(a) and 7(c)). Since there is a good agreement between the FEA results and experimental data, two further FEAs were undertaken. The nominal capacity of the steel material (blind bolts, nut, washer and steel tube) is used in the FEAs, as would be done in a design situation. First, the relative contributions to the tensile resistance from the bearing of the headed anchor on the infill concrete (the so-called concrete contribution ) and from the bearing of the washer on the inside of the tube wall (the so-called tube wall contribution ) was calculated at various load levels. Second, results of parametric studies on concrete strengths and embedment depths were obtained. 5. Concrete and tube wall contribution using FEAs The purpose of this analysis was to determine practical combinations of the thickness of the steel tube and the embedment depth of the HABB that were sufficient to achieve the ultimate capacity of HABB. The washer bearing on the steel tube wall and the headed anchor bearing on the concrete combine to resist the pull-out force applied to the HABB. In the beginning, the concrete contribution is higher than the steel tube wall contribution. As cracks develop in the concrete, the stiffness of the concrete cone decreases, resulting in a decrease in the concrete contribution. Using FEAs, the tube wall contribution was measured in case N1, and the concrete contribution was investigated by eliminating the washer from the finiteelement assembly. Figure 8 shows both the concrete contribution (case N2_nw, where nw indicates no washer ) and the tube wall contribution (case N1). Figure 8 also shows that case N2 is equal to the summation of case N1 and case N2_nw (the summation between case N1 and N2_nw is shown in Figure 8 as sum ) when the pull-out force is lower than 7% of the nominal ultimate capacity. Applying a larger force than this caused yielding of the blind bolt in the combined case N2, T6_D16_N1_F T6_D16_N2_F T6_D16_N2_F_nw Sum (a) T8_D2_N1_F T8_D2_N2_F 4 T8_D2_N2_F_nw Sum (d) Sum (b) T6_D2_N1_F T6_D2_N2_F T6_D2_N2_F_nw T1_D2_N1_F T1_D2_N2_F 4 T1_D2_N2_F_nw Sum (e) T8_D16_N1_F T8_D16_N2_F 4 T8_D16_N2_F_nw Sum (c) Figure 8. Concrete and tube wall contribution: (a) T6_D16; (b) T6_D2; (c) T8_D16; (d) T8_D2; (e) T1_D2 754 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved.

9 Volume 168 Issue SB1 whereas the blind bolt in the individual case N1 and case N2_nw would still be in the elastic state, and hence inaccuracy would occur if direct summation of case N1 and case N2_nw were to be performed for loads exceeding 7% of the nominal ultimate capacity. Since it is assumed that the force in a blind bolt in a momentresisting connection will not be allowed to exceed 6% of the nominal ultimate capacity of the bolt, further analysis is only performed up to 7% of the ultimate nominal capacity. The concrete contribution is expressed as a percentage of the overall pull-out resistance in Figure 9 for applied loads of 1%, 3%, 6% and 7% of the ultimate nominal capacity. It is shown that for cases in which the embedment depth is constant at 1 mm, either decreasing the CHS thickness or enlarging the blind bolt diameter will increase the concrete contribution when in the range of 1 7% of the nominal ultimate capacity. The concrete contribution generally decreases as the pull-out force increases, as has been discussed above. 6. Parametric studies using FEAs Parametric studies on the concrete strength and embedment depth are reported in this section. These parametric studies were undertaken on specimens with various practical D/t ratios (from 32 4 to 54) and BB or HABB diameters (16 and 2 mm). Concrete compressive strengths of 2, 32, 4 and 48 MPa were used for the specimens with an embedment depth of 1 mm. Moreover, embedment depths of, 5, 75, 1 and 125 mm were investigated in the specimens with a compressive strength equal to 48 MPa. Concrete contribution: % T6_D16_N2_F T6_D2_N2_F T8_D16_N2_F T8_D2_N2_F T1_D2_N2_F Percentage of the nominal ultimate capacity, F u : % Figure 9. Percentage of the concrete contribution to the pull-out resistance at 1%, 3%, 6% and 7% of the nominal ultimate strength 6.1 Parametric study on the concrete strength Figure 1 shows that the concrete strength has little influence on the stiffness and strength when the compressive strength of the concrete is between 4 and 48 MPa. At 6% of the ultimate nominal capacity (shown as 6F u in Figure 1), the stiffness observed in the specimens with compressive strengths of 4 and 48 MPa is similar. Given the poor behaviour exhibited in Figure 1(b) and 1(d), it is recommended that concrete with a compressive strength equal to 2 MPa is not used, since the ultimate nominal capacity of the blind bolt cannot be reached. Figure 1 shows that the thickness of the steel tube (D/t ratio) is also very important: as the blind bolt diameter is increased from 16 to 2 mm, either the D/t ratio should be decreased (Figure 1(d)) or the concrete strength should be increased (Figure 1(b)) to ensure that the blind bolt capacity can be reached by the combined contribution from the concrete and steel tube. In general, as the concrete compressive strength increases, the stiffness also increases. 6.2 Parametric study on the embedment depth The full stiffness and capacity is defined in this paper as the highest stiffness (defined as the secant stiffness at 6% of the nominal ultimate tensile strength of the bolt) and capacity that can be achieved for a certain D/t ratio, HABB diameter, and for a concrete infill with an average compressive strength of at least 48 MPa; there is a minimum embedment depth needed to achieve this. The highest capacity corresponds to the ultimate tensile strength of the HABB. It is shown by all the plots in Figure 11 that all specimens with an embedment depth of 5 mm exhibit premature stiffness degradation. This occurs because the concrete cone forms and bears on the tube wall. As the thickness of tube wall is increased (Figure 11(a) to 11(c) and Figure 11(b) to 11(d)), the stiffness does not decrease as rapidly because the bearing contribution of the tube wall increases. An embedment depth of 75 mm (4 7 times the blind bolt diameter) is sufficient to develop the full stiffness and capacity only for a specimen with a 16 mm blind bolt diameter and at least an 8 mm tube wall thickness, as for specimen T8_D16_N2_F (Table 4 and Figure 11(c)). With an increase in the diameter of the blind bolt from 16 to 2 mm (T8_D2_N2_F), a minimum embedment depth of 1 mm (five times the blind bolt diameter) is needed to achieve the full stiffness and capacity of the HABBs. As mentioned previously, the stiffness at 6% of the capacity is the most important factor for design purposes. Figure 11 shows that only specimens T6_D2_N2_F (Figure 11(b)) and T8_D2_N2_F (Figure 11(d)) need a 1 mm minimum embedment depth to achieve full stiffness at 6% of the ultimate capacity. Table 4 summarises the minimum embedment depth that should be used for the given configurations in order to achieve full stiffness and capacity, provided that the average concrete compressive strength, f cm, is greater than or equal to 4 MPa. In general, as long as the full capacity is reached, the full stiffness is also reached. However, even if the full stiffness Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 755

10 Volume 168 Issue SB F u T6_D16_N2_F_2 MPa T6_D16_N2_F_32 MPa T6_D16_N2_F_4 MPa T6_D16_N2_F_48 MPa F u T6_D2_N2_F_2 MPa T6_D2_N2_F_32 MPa T6_D2_N2_F_4 MPa T6_D2_N2_F_48 MPa (a) (b) F u T8_D16_N2_F_2 MPa T8_D16_N2_F_32 MPa T8_D16_N2_F_4 MPa T8_D16_N2_F_48 MPa F u T8_D2_N2_F_2 MPa T8_D2_N2_F_32 MPa T8_D2_N2_F_4 MPa T8_D2_N2_F_48 MPa (c) (d) F u T1_D2_N2_F_2 MPa T1_D2_N2_F_32 MPa T1_D2_N2_F_4 MPa T1_D2_N2_F_48 MPa (e) Figure 1. Parametric study on the concrete strength with a 1 mm embedment depth: (a) T6_D16_N2; (b) T6_D2_N2; (c) T8_D16_N2; (d) T8_D2_N2; (e) T1_D2_N2 at 6% of the ultimate capacity is reached, the full capacity may not be reached. 7. Conclusion The results of pull-out tests on BBs and HABBs in CFCHS columns are reported here and compared with results from FEA. Specimens with D/t ratios from 32 4 to 54 were used. Several parameters such as the tube thickness, HABB diameter and embedment depth were varied both in the experiments and FEA. Since both the experimental and FEA results were in good agreement, the concrete contribution and the steel tube contribution to the load-carrying capacity were determined using FEA. Parametric studies covering different concrete strengths and varying embedment depths were also performed. The main conclusions are summarised below. & The stiffness of the bolted connection in tension depends on the D/t ratio of the tube and the diameter of the blind bolt for the case of blind bolts without embedded heads (case N1). On the other hand, for the case of HABBs (case N2), the stiffness depends on the concrete strength, the D/t ratio of the tube, the diameter of the bolt and the embedment depth. 756 Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved.

11 Volume 168 Issue SB F u T6_D16_N1_F T6_D16_N2_F_5 mm T6_D16_N2_F_75 mm T6_D16_N2_F_1 mm T6_D16_N2_F_125 mm F u T6_D2_N1_F T6_D2_N2_F_5 mm T6_D2_N2_F_75 mm T6_D2_N2_F_1 mm T6_D2_N2_F_125 mm (a) (b) F u T8_D16_N1_F T8_D16_N2_F_5 mm T8_D16_N2_F_75 mm T8_D16_N2_F_1 mm T8_D16_N2_F_125 mm F u T8_D2_N1_F T8_D2_N2_F_5 mm T8_D2_N2_F_75 mm T8_D2_N2_F_1 mm T8_D2_N2_F_125 mm (c) (d) F u T1_D2_N1_F T1_D2_N2_F_5 mm T1_D2_N2_F_75 mm T1_D2_N2_F_1 mm T1_D2_N2_F_125 mm Figure 11. Parametric study on different HABB embedment depths for a concrete strength of 48 MPa: (a) T6_D16; (b) T6_D2; (c) T8_D16; (d) T8_D2; (e) T1_D2 (e) & & For N1, increasing the thickness of the tube wall (decreasing the D/t ratio) is more effective in increasing both the stiffness and strength than enlarging the diameter of the blind bolt. On the other hand, N2 specimens exhibit the reverse of this. Moreover, by modifying the blind bolt by adding a headed anchor (case N2), the ultimate nominal capacity of the blind bolt can be achieved, provided the embedment depth and concrete strength are adequate (see the final two points below). At low loads, the stiffness of the headed anchor is governed by the bearing action of the head on concrete. As the load level increases and micro-cracks start to form in the & concrete, the stiffness becomes dependent on the bearing of the bolt head on the tube wall. Accordingly, the balance of load sharing between the concrete and tube wall changes as the load increases on the bolt. It is recommended that, for specimens with an embedment depth of 1 mm, concrete with a compressive strength less than or equal to 2 MPa is not used, because the ultimate nominal capacity of the blind bolt cannot be reached. Concrete with a compressive strength greater than or equal to 4 MPa is ideal. In general, as the concrete compressive strength increases above 4 MPa, the stiffness increases only slightly. Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 757

12 Volume 168 Issue SB1 Specimen Blind bolt diameter: mm D/t ratio Minimum embedment depth to achieve full stiffness and capacity: mm Minimum embedment depth to achieve full stiffness at 6% of the ultimate capacity: mm Embedment depth/ HABB diameter to achieve full stiffness and capacity Embedment depth/ HABB diameter to achieve full stiffness at 6% of the ultimate capacity T6_D16_N2_F T6_D2_N2_F T8_D16_N2_F T8_D2_N2_F T1_D2_N2_F Table 4. Minimum embedment depths for specific D/t ratios and HABB diameters for f cm 4 MPa & It is important to ensure that the ultimate tensile strength of the HABB can be reached by the combined contribution from the anchorage of the head into the concrete and the bearing of the bolt head and washer onto the wall of the steel tube. Minimum embedment depths to achieve the full stiffness and capacity have been determined for a combination of various tube D/t ratios and concrete strength of greater than or equal to 4 MPa, as given in Table 4. Acknowledgements The authors would like to acknowledge the Australian Research Council (ARC) and collaborating organisations, Orrcon Steel and Ajax Engineered Fasteners, for supporting this research project, which is funded by Linkage grant LP REFERENCES ABAQUS (211) ABAQUS Documentation. Dassault Systèmes Simulia, Providence, RI, USA. ACI (211) ACI 318M: Building code requirements for structural concrete (ACI 318M-11) and commentary. American Concrete Institute, Farmington Hills, MI, USA. Agheshlui H (214) Anchored Blind Bolted Connections within Concrete Filled Square Steel Hollow Sections. PhD thesis, University of Melbourne, Melbourne, Australia. AISC (21) Specification for structural steel buildings. American Institute of Steel Construction, Chicago, IL, USA. Ajax Engineered Fasteners (22) ONESIDE brochure B-N12 Data Sheet. Ajax Engineered Fasteners, Victoria, Australia. Fernando S (25) Joint Design using ONESIDETM Structural Fastener. Ajax Engineered Fasteners, Victoria, Australia, Technical note AFI/3/12. France J, Davison B and Kirby P (1999) Strength and rotational stiffness of simple connections to tubular columns using flowdrill connectors. Journal of Constructional Steel Research 5(1): Guezouli S and Lachal A (212) Numerical analysis of frictional contact effects in push-out tests. Engineering Structures 4: Han LH and Li W (21) Seismic performance of CFST column to steel beam joint with RC slab: experiments. Journal of Constructional Steel Research 66(11): Huck International (199) Industrial Fastening Systems. Huck International, Tucson, AZ, USA. Lee J, Goldsworthy HM and Gad EF (21) Blind-bolted T-stub connections to unfilled hollow section columns in low-rise structures. Journal of Constructional Steel Research 66(8 9): Lee J, Goldsworthy HM and Gad EF (211a) Blind bolted moment connection to sides of hollow section columns. Journal of Constructional Steel Research 67(12): Lee J, Goldsworthy HM and Gad EF (211b) Blind bolted moment connection to unfilled hollow section column using extended T-stub with back face support. Engineering Structures 33(5): Lindapter International (1995) Type HB Hollo-bolt for Blind Connection to Structural Steel and Structural Tubes. Lindapter International, Bradford, UK. Mahmood M, Tizani W and Sansour C (214) Effect of tube thickness on the face bending for blind-bolted connection to concrete filled tubular structures. International Journal of Civil, Architectural, Structural and Construction Engineering 8(9): Mander JB, Priestley MN and Park R (1988) Theoretical stress strain model for confined concrete. Journal of Structural Engineering 114(8): Pitrakkos T and Tizani W (213) Experimental behaviour of a novel anchored blind-bolt in tension. Engineering Structures 49: Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved.

13 Volume 168 Issue SB1 Standards Association of Australia (1991) AS1163: Structural steel hollow sections. Standards Australia, Homebush, Australia. Standards Australia (1998) AS 41: Steel structures. Standards Australia, Sydney, Australia. Standards Australia (2) AS 4291: Mechanical properties of fasteners made of carbon steel and alloy steel. Standards Australia, Strathfield, Australia. Standards Australia (29) AS 36: Concrete structures. Standards Australia, Sydney, Australia. Standards Australia (21) AS 3972: General purpose and blended cements. Standards Australia, Sydney, Australia. Yao H, Goldsworthy HM and Gad EF (28) Experimental and numerical investigation of the tensile behaviour of blind-bolted T-stub connections to concrete-filled circular column. Journal of Structural Engineering 134(2): Yao H, Goldsworthy HM, Gad EF et al. (211) Experimental study on modified blind bolts anchored in concrete-filled steel tubular columns. Australian Earthquake Engineering Society Conference, Barossa Valley, Australia. WHAT DO YOU THINK? To discuss this paper, please up to 5 words to the editor at journals@ice.org.uk. Your contribution will be forwarded to the author(s) for a reply and, if considered appropriate by the editorial panel, will be published as discussion in a future issue of the journal. Proceedings journals rely entirely on contributions sent in by civil engineering professionals, academics and students. Papers should be 2 5 words long (briefing papers should be 1 2 words long), with adequate illustrations and references. You can submit your paper online via where you will also find detailed author guidelines. Downloaded by [ University of Melbourne] on [11/1/15]. Copyright ICE Publishing, all rights reserved. 759