Effect of bolt gauge distance on the behaviour of anchored blind bolted connection to concrete filled tubular structures

Transcription

1 Tubular Structures XV Batista, Vellasco & Lima (eds) 2015 Taylor & Francis Group, London, ISBN Effect of bolt gauge distance on the behaviour of anchored blind bolted connection to concrete filled tubular structures M. Mahmood Department of Civil Engineering, The University of Nottingham, UK Department of Civil Engineering, University of Diyala, Iraq W. Tizani & C. Sansour Department of Civil Engineering, The University of Nottingham, UK ABSTRACT: This study focuses on investigating the effect of bolt gauge distance on the column face bending behaviour for anchored blinded bolting connections to concrete filled tubular columns. Full scale experimental tests were conducted for six samples and ABAQUS 6.13 was employed to perform a parametric study to extend the experimental data. Results show that, increasing the bolt gauge by 75% and 125% can improve the column face bending strength by 26% and 58% respectively and the initial stiffness by 22% and 44% respectively. At the ultimate resistance of the component and for all the bolt gauges that were investigated in this study, complete yielding of the tubular corners was recorded. This finding contradicts with exciting knowledge which assumes that the yield lines could extend only to the centreline of the SHS wall. 1 INTRODUCTION Concrete filled tubular columns have high strength, high stiffness and high ductility. The concrete and the tubular section work as a complementary part to each other. The concrete provides the support to the tubular walls and improve their buckling strength. The tubular structures provide the confinement to the concrete and increase its strength. In addition, the tubular section can act as a formwork during the construction period, which can reduce the cost and construction time (Ellobody andyoung, 2006). However, performing the connection to this type of construction represents a challenge due to the difficulty of access to the inside of the tubular sections. Lindapter company patented a practical blind bolting system called HolloBolt (HB) (Figure 1,a). It s installation needs only inserting it in the pre drilled fixture and steel work and tightening it to open the sleeves inside the tubular column and perform the connection. HB has a 40% higher shear strength than the standard bolt (Occhi, 1995). However, it has a low stiffness compared with standard bolt under tensile load (Barnett et al., 2000). To overcome the problem of low stiffness, Tizani and Ridley-Ellis (2003) developed a novel Extended anchored HolloBolt (termed EHB) (Figure 1, b) fastener. The idea of the development is based on increasing the shank length and attaching an anchoring nut at the shank s end. Embedding the extended shank length and the nut inside the concrete improves the behaviour of the fastener through anchoring the EHB in the concrete. The extended shank length and nut size are limited to the size of the tubular section and the diameter of the bolt hole. Pitrakkos and Tizani (2013) carried out an extensive experimental program to investigate the tensile behaviour of EHB under pure tension. The study confirmed that the tensile capacity of the EHB component is controlled by its internal bolt tensile capacity. The work presented in this paper is part of a series of studies, which are investigating the behaviour of the different components of EHB connection under variety of loading conditions to provide the required knowledge for producing design guidance for EHB connections. In this study, the focus is set on investigating the effect of bolt gauge distance on the bending behaviour of the column face component for EHB connections. 2 EXPERIMENTAL TESTS 2.1 Test specimens A series of experimental tests were conducted to investigate the effect of bolt gauge (g) on the column face bending behaviour. All the tests were performed using SHS and M16 EHB (dummy EHB). The shape and the dimensions of the dummy EHB are exactly similar to the installed EHB, and it was manufactured from high strength steel alloy (Elamin, 2013) to limit the failure at the column face. The minimum and maximum values of the bolt gauge are controlled by the geometry of the EHB and the tubular section. 87

2 Figure 2. Maximum and minimum bolt gauge distance. Figure 1. HolloBolt (HB) and Extended HolloBolt (EHB). As well as providing a suitable distance for concrete placement, which is not less than 20 mm according to Eurocode (CEN, 2004b). The maximum opening width of the sleeves was calculated theoretically and it was found equal to 46.6 mm (Figure 2). Therefore, the minimum bolt gauge (g min ) required to satisfy the previous conditions was calculated as follows: Similarly, the minimum edge distance between the sleeves edges and the inside wall of the tubular section should not be less than 20 mm.therefore, Equation (1) was suggested to calculate the maximum bolt gauge (g max ). where b = tubular section width (mm) and t = tubular section plate thickness (mm). The experimental programme was carried out by testing a pair of identical samples for each bolt gauge to confirm the reliability of the results. Six samples were tested: two for bolt gauge of 80 mm (G80-1 & G80-2), two for bolt gauge140 mm (G140-1 & G140-2) and two for bolt gauge of 180 mm (G180-1 & G180-2). Apart from the bolt gauge, all the geometrical and material properties and the test procedure were similar. The concrete mix was designed to have a compressive strength of 40 N/mm 2 on the day of the test. The anchored length of the EHB was 80 mm and all the samples were cut from one batch of tubular section. The actual section dimensions and the holes size for all the tested samples are listed in Table 1. These values were used in the validation of the finite element model. Table 1. Actual section dimensions and holes size for the tested samples. G80-1 G80-2 G140-1 G140-2 G180-1 G180-2 Section width (b) mm Plate thickness (t) mm Corner thickness (t c ) mm Hole diameter (d) mm 2.2 Test setup and instrumentations The test setup is presented in Figure 3. The test rig was designed to apply pull-out load on the EHBs. The loading was displacement controlled with a rate of mm/sec. During the test, a video gauge system was employed to record the column face displacement, bolt slip and the sample movement. This system can measure a wider range of displacement compared with the linear potential meter and it was used to provide accurate results in similar studies (Mahmood et al., 2014). The sample movement was recorded to eliminate it from the column face displacement and the bolt slip. Recording the bolt slip requires attaching some fittings to the sample. A 300 mm length M16 threaded rode was attached to the end of the EHB inside the sample and it passed through a hole in the back of the sample. This rod was covered by 20 mm diameter steel tube to protect it from concrete during the casting (Figure 4). A Digital Image Correlation (DIC) system was used to monitor the strain distribution in the column face. 88

3 Figure 3. Test arrangement. Figure 4. Sample setup. 2.3 Test results The specified cube concrete compressive strength in the day of the test was 39.3, 40.1 and 40 N/mm 2 for samples with bolt gauge of 80, 140 and 180 mm respectively. The effect of variation in the concrete strength was assumed to be neglected since the maximum difference is 0.8 N/mm 2. The force displacement curves for all the tested samples are presented Figure 5. Samples G80-1 and G80-2 show early nonlinear behaviour at about 66% of the yield load, whereas the linear behaviour for samples G14-1, G140-2, G180-1 and G180-2 extends up to about 80% of the ultimate strength. The early nonlinear behaviour for components with small bolt gauge might be due to the limited role of anchoring. In this study, the ultimate strength of the connection represents the peak point before the strength starts dropping. Once the ultimate strength was reached, there were some concrete crushing sounds coming from the samples. These sounds were synchronised with a drop in the strength, and this scenario was repeated for all the tested samples. However, the rate of the drop was increased with the use of large bolt gauge. A possible explanation for this behaviour is the use of small bolt gauge results in a stress concentration at the concrete between the bolts (the gauge distance). This limits the advantages of anchoring the EHB in the concrete through the concrete failure and losing the composite action with the bolts at early stage. However, with the use of large bolt gauge, the stress distributed over a wider concrete area. Figure 5. Load versus column face displacement curves. This results in delaying the concrete failure and maintaining the connection strength to much higher than that for the small bolt gauge. Therefore, the concrete failure causes sudden drop in the strength. After this drop, the connection strength starts increasing due to the membrane action in the column face. Results indicate that, increasing the bolt gauge by 75% and 125% could improve the ultimate strength of the connection up to 26% and 58% respectively and the initial stiffness up to 22% and 44% respectively. The previous results suggest that, increasing the bolt gauge can improve the column face bending strength 89

4 Figure 7. Bolt gauge effect on bolt movement. Figure 6. Table 2. Sample Column face displacement and bolt slip. Pull-out displacements. Column face displacement (mm) G G G G G G Figure 8. Strain outline sample G more than the stiffness. The improvement comes from changing the loading distribution on the component by varying the bolt gauge. As increasing the bolt gauge, results in decreasing the shear span of the applied load which requires more work to bend the column face component. Figure 6 shows comparison between the column face displacement and the bolt slip for all the tested samples, and the values of the pulling-out displacements are summarized in Table 2. The pull-out point is identified when the column face displacement becomes less than the bolt slip at the same load value. It is noticeable that for larger gauge distances, the bolts tend to pull-out earlier (at a lower face displacement) than those with relatively smaller gauge distance. This may be attributed to the different in the higher load and also when the bolt starts pulling-out, it begins pushing the hole sides outward. With low values of bolt gauge, the distance between the bolts is small and each bolt is trying to push the column face plate towards the another bolt. This increases the pressure of the holes on the sleeves and results in holding the bolts in the holes for higher displacement (Figure 7a). However, with large bolt gauge there is a long distance between the bolts so that the hole can expand easier to allow the bolt pulling-out (Figure 7b). The DIC results show that the first yield was at the hole sides at 75% of the ultimate strength.then it starts propagating in all directions. At the ultimate resistance of the connection, the yield pattern has a shape of half circle with a radius extend from the hole center to corner of the tubular section. Then, propagation of the yield was mainly in the column face. In contrast to the previous understanding which assumes that the yielding could only extend to the centreline of the tubular corners, this finding confirmed that the whole corner of the tubular section yields at the ultimate load. The yield outline is presented in Figure 8, only half of sample was monitored due the limited access to the column face during the test. 90

5 and dimensions similar to the EHB was created at the bolt location. The concrete elastic behaviour was simulated by the Young s modulus of elasticity (E c ) and Poison s ratio. E c was calculated using equation (2) (CEN, 2004a) and the poison s ratio was considered as 0.2. Figure 9. FE details. 3 FINITE ELEMENT MODELLING AND ANALYSIS ABAQUS 6.13 (Dassault, 2013) was used to perform Finite Element (FE) analysis. The FE model consists of three parts: tubular section, concrete and the EHBs. To reduce the modelling and computational cost, the advantage of symmetry was considered and only quarter of the sample was modelled. The material and geometrical properties and the boundary conditions were considered to be similar to the real samples. The details of the FE model are presented in 3.1. All the model components were modelled using (C3D8) element. This element is suitable for simulating the complete nonlinear behaviour including contact and geometrical nonlinearities (Abaqus, 2013). The contact between the three parts of the model was modelled using surface based and pair algorithm. Both normal and tangential contacts are existing in the model. The coefficient of friction between steel and concrete was considered as 0.25 (Elr y and Azizinamini, 2001, Hu et al., 2003) and between steel and steel was considered as 0.45.(Wang, 2012). The analysis was performed by applying a static uniform displacement at the EHB similar to the experimental test. The increments were calculated using RIKS method. This method is powerful in predicting the nonlinear behaviour of the structures, because it has the ability for simulating the softening part in the behaviour. 3.1 Modelling of tubular section The actual dimensions that are summarised in Table 1 were used to create the geometry of the tubular section in the model. The measured material properties reported by Mahmood et al. (2014) were used in this study. The experimental measured yield strength (f y ) was 406 N/mm 2, the ultimate strength (f u )was 537 N/mm 2, theyoung s modulus of elasticity (E s )was N/mm 2 and the Poison s ratio was Modelling of concrete The concrete geometry was modelled using the interior dimensions of the tubular section. A hole with a shape where E c = concrete Young s modulus (N/mm 2 ) and f cm = mean value of concrete cylinder compressive strength (N/mm 2 ). The plastic behaviour of the concrete was simulated using the Concrete Damage Plasticity (CDP) model. The failure mechanisms for this model are the tensile cracking and compression crushing. The concrete compression stress-strain curve was predicted using the Euro code multi-linear model for nonlinear structural analysis (CEN, 2004a). The model assumes that the concrete has a linear behaviour until 0.4f cm. Then, then the behaviour was predicted using the following equations. where f c = concrete compressive stress at any point on the stress-strain curve (N/mm 2 ), ε c1 = concrete compressive strain at the maximum stress (f cm ), ε c2 = concrete compressive strain at the end of stressstrain curve and ε c = compressive strain in the concrete. The following plasticity parameters were used: dilation angle is 35, eccentricity is 0.1, the ratio of initial equibiaxial compressive yield stress to initial uniaxial compressive yield stress (σ bo /σ co ) is 1.16, yield shape parameter (kc) is 1 and viscosity parameter (µ o ) is 0. The tensile behaviour of concrete was simulated using a bilinear model with ultimate strength of 0.1f cm. 3.3 Modelling of EHB The geometry of the EHBs is exactly similar to dummy EHB. They are assumed to have high strength, therefore, they were modelled as elastic material using Young s modulus of N/mm 2 and poison s ratio of 0.3 (Elamin, 2013). 91

6 Table 3. strength. Experimental and numerical ultimate column face Ultimate load (kn) Sample Experimental Numerical Difference % G G G G G G Figure 10. Load strain curves: experimental versus FE at bolt hole, sample G VERIFICATION OF FINITE ELEMENT MODEL A comparison between the experimental and FE analysis results was carried out to verify the FE model. The comparison was considered the following parameters: ultimate column face strength strain values at the bolt hole and the load versus the column face displacement behaviour. The maximum difference in the ultimate column face strength was 9% and it was recorded for sample G The experimental and numerical ultimate column face strength is listed in Table 3. Figure 10 presents a comparison between the experimental and numerical results for the stain values at the bolt hole. Good correlation between the both results was achieved. The numerical and experimental load versus column face displacement for all samples is plotted in Figure 11. The numerical results agreed very well with the experimental results up to the ultimate column face strength. Then, the FE model shows stiffer behaviour than experimental results due to the limited damage in the concrete model. 5 PARAMETRIC STUDY A parametric study was conducted using the validated FE model. Sample G140-1 was used as a base Figure 11. versus FE. Load-column face displacement, experimental sample. Eleven models were analysed with different values of bolt gauge, ranged from 80 mm to 180 mm. The applied load versus the column face displacement curves are plotted in Figure 12. Although the bolt gauge distance was increased by constant value (10 mm), the amount of improvement in the ultimate column face strength was increased with the use of higher bolt gauge. For instance, increasing the bolt gauge from 80 mm to 90 mm improved the column face strength by %2, whereas increasing the bolt 92

7 ACKNOWLEDGMENT The authors wish to acknowledge TATA Steel and Lindapter International for supporting work. Gratitude is expressed to Mr Trevor Mustard, of TATA Steel, and Mr Neil Gill, of Lindapter International. The first author would like to thank the higher committee for education development in Iraq for providing the chance to perform this research. REFERENCES Figure 12. Effect of bolt gauge on the column face bending behaviour. gauge from 170 mm to 180 mm enhanced the column face strength by %8. Also the connection could reach its ultimate strength at lower column face deformation with the use of large bolt gauge. For example the connection with a bolt gauge of 80 mm reaches its ultimate strength at column face displacement of mm, while the connection with a bolt gauge of 180 mm reaches its ultimate strength at column face displacement of mm. 6 CONCLUSIONS This paper presented an experimental and numerical investigation of the effect of bolt gauge distance on the bending behaviour of the column face component for a novel anchored blind bolted connection to concrete filled tubular columns.the behaviour of all the investigated samples was approximately linear up to the ultimate resistance of the column face. The sharpness of the drop in the strength after the ultimate column face strength was increased with the use of large bolt gauge. The bolt gauge distance has a clear influence at the strength and stiffness of the column face. However, the use of small bolt gauge could limit the advantages of using the anchored blind bolting system. The ultimate strength of the connection could be reached with less deformation and the bolt starts bulling out at lower column face deformation with the use of large bolt gauge. The yield outline was mainly in the column face and tubular corners. Complete yielding of the corners at the ultimate resistance was recorded, contrary to standard understanding that the yielding could only extend to the centreline of the corners. Abaqus, I ABAQUS Analysis User s Manual: Volume IV: Elements. Abaqus, Inc. Dassault Systèmes. Barnett, T., Tizani, W. & Nethercot, D Blind Bolted Moment Resisting Connections to Structural Hollow Sections. Connections in Steel Structures IV: Steel Connections in the New Milennium, CEN 2004a. Eurocode 2: Design of Concrete Structures: Part 1-1: General Rules and Rules for Buildings. EN British Standards Institution. CEN 2004b. Eurocode 4: Design of Composite Steel and Concrete Structures: Part 1-1: General Rules and Rules for Buildings. EN CEN. Dassault Abaqus v [Software]. Dassault Systèmes Simulia Corp. Providence, RI: Dassault Systèmes Simulia Corp. Elamin, A The Face Bending Behaviour of Blind- Bolted Connections to Concrete-Filled Hollow Sections. PhD Thesis, University of Nottingham. Ellobody, E. & Young, B Nonlinear Analysis of Concrete-Filled Steel SHS and RHS Columns. Thinwalled structures, 44, Elr y, A. & Azizinamini, A Design Provisions for Connections between Steel Beams and Concrete Filled Tube Columns. Journal of Constructional Steel Research, 57, Hu, H. T., Huang, C. S., Wu, M. H. & Wu, Y. M Nonlinear Analysis of Axially Loaded Concrete-Filled Tube Columns with Confinement Effect. Journal of Structural Engineering, 129, Mahmood, M., Tizani, W. & Sansour, C 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, Occhi, F Hollow Section Connections Using (Hollofast) Hollobolt Expansion Bolting. Second Interim Report 6G-16/95. Sidercad, Italy. Pitrakkos, T. & Tizani, W Experimental Behaviour of a Novel Anchored Blind-Bolt in Tension. Engineering Structures, 49, Tizani, W. & Ridley-Ellis, D. The Performance of a New Blind-Bolt for Moment-Resisting Connection. TUBULAR STRUCTURES-INTERNATIONAL SYMPOSIUM-, Wang, Z Hysteretic Response of an Innovative Blind bolted Endplate Connection to Concrete Filled Tubular Columns. PhD Thesis, University of Nottingham. 93