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Experimental and numerical analysis of a bolted connection in steel transmission towers
Journal of Constructional Steel Research 121 (2016) 253–260
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Journal of Constructional Steel Research
Experimental and numerical analysis of a bolted connection in steel transmission towers Eray Baran a,⁎, Tolga Akis b, Gokmen Sen c, Ammar Draisawi b a b c
Department of Civil Engineering, Middle East Technical University, Ankara 06800, Turkey Department of Civil Engineering, Atilim University, Ankara 06836, Turkey MITENG, Ankara 06800, Turkey
a r t i c l e
i n f o
Article history: Received 9 July 2015 Received in revised form 21 January 2016 Accepted 12 February 2016 Available online xxxx Keywords: Lattice transmission tower Topic: Bolted connection Splice connection Steel angle Bolt slip Finite element analysis Tension member
a b s t r a c t This paper presents an integrated numerical and experimental study on a bolted splice connection used in main legs of steel lattice transmission towers. At specific locations, where the number of angle sections in built-up cross section of main leg members changes, the complex geometry around the connection region results in eccentricities in the load path and indirect load transfer. Such complex configurations and uncertainties in the load path have led to overdesigned connections with increased number of bolts and redundant connection reinforcing members. The current study was conducted in an attempt to gain a better understanding of the load-flow mechanism at this specific location where the cross section of main leg members changes. The experimental part included tensile load testing of six specimens with different connection details. The main parameters used in the testing program were the number of bolts used in the connection as well as the presence of connection reinforcement angles and tie plate. For all connection configurations studied, the failure occurred due to net section fracture of upper main member angle near leading bolt holes. The calculated load capacity based on the measured material strength closely predicted the measured load capacity of specimens. The experimentally determined response of each connection configuration was better predicted by the FE model that incorporates bolt slip as compared to the model that assumes no slip. The experimental and numerical results also indicate that major differences among the investigated connection details do not cause any appreciable difference in behavior under tensile loading. © 2016 Elsevier Ltd. All rights reserved.
1. Introduction Steel lattice towers have been extensively used worldwide as an important component of power transmission networks. These towers are usually constructed using equal leg angle section members and bearing type bolted connections. Depending on the magnitude of conductor and environmental loads as well as the geometric constraints of the transmission line, the axial force in main members at lower parts of the tower could get so large that built-up sections have to be used at these locations. These built-up sections are usually made of two, three, or four steel angle sections connected corner-to-corner, depending on the level of axial force in the member. At higher locations along the height of tower, axial force level in main legs decreases and this allows a reduction in the cross-sectional area of these members. Such reduction in the cross-sectional area of main legs is usually achieved by reducing the number of steel angle sections in the built-up cross section. At the specific locations, where the number of angle sections in the builtup cross section changes, the complex geometry around the connection region results in eccentricities in the load path and indirect load transfer ⁎ Corresponding author. E-mail address: [email protected] (E. Baran).
http://dx.doi.org/10.1016/j.jcsr.2016.02.009 0143-974X/© 2016 Elsevier Ltd. All rights reserved.
between several components. Such complex configurations and the uncertainties in the load path have led to overdesigned connections with increased number of bolts and redundant connection reinforcing members. The current study was conducted in an attempt to gain a better understanding of the load-flow mechanism at this specific location where the cross section of main leg members changes. The overall goal was to simplify the connection geometry by identifying the redundant members that can be eliminated from the connection without significantly affecting the structural behavior. An experimental study integrated with numerical analyses was conducted in order to achieve this goal. Numerical analysis of steel lattice towers has recently been performed by many researchers [1–4]. These numerical analyses were conducted using finite element models constructed using either one-dimensional beam–column elements or two-dimensional shell and plate elements. Eccentricity of connections and flexibility of joints were also considered in some of the studies. In majority of the cases, results obtained from the models were compared to the observed response from full-scale proof-load tests or smaller scale laboratory tests of the towers. These studies indicate that the maximum load capacity and the distribution of member failures can be predicted with a fair accuracy using numerical models. Based on the analysis of transmission
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line tower failures during prototype load testing, the possible reasons for tower failures are given as incorrect design assumptions, improper detailing, material defects, fabrication errors, force fitting during erection, and variation in bolt strength [1]. Several studies have also focused on the effect of bolt slip on nonlinear behavior of lattice towers [5–7]. Results from these studies indicated that the joint slip would be expected to affect the tower displacement, but have no major influence on the failure mode and ultimate strength of the tower. Another current area of research related with overhead transmission line towers is on strengthening of existing towers. Various strengthening methods to improve the load capacity of existing lattice towers so that they can tolerate the increase in conductor loads have recently been studied both experimentally and numerically [8–12]. These studies revealed that upgrading of existing towers could be achieved by providing additional bracing for members or by increasing the total cross-sectional area of individual members by reinforcement. Experimental and numerical studies on bolted splice connections of different types of metallic members are also available in the current literature [13–18].These studies revealed information regarding many aspects of bolted connections that can be useful in design of various bolted connections in steel lattice towers. The main phenomena investigated with these studies are the influence of connection geometry on the extent of shear lag effect on steel angles, the influence of the amount of edge distance as well as the level of bolt pretension on the behavior of plate connections, and the influence of connection eccentricity on the behavior of T-section tension members. 2. Experimental program The experimental part of the study included tensile load testing of six specimens with different connection details. The main parameters used in the testing program were the number of bolts used in the connection as well as the presence of connection reinforcement angles and tie plate. The specimens were fabricated at the production facility of a local manufacturer of transmission line towers, and were tested at Atilim University Structural Mechanics Laboratory. Fig. 1. Prototype design of 75 m tall 500 kV capacity tower.
2.1. Test specimens A prototype latticed transmission tower was designed, and the splice connection at main legs of this prototype structure was used as a basis for the test specimens. General geometry of the prototype tower and the location of the investigated splice connection along the length of the tower are shown in Fig. 1. The loads considered during design of the 75 m tall 500 kV prototype tower included wind, ice, equipment, conductor, broken wire, and earthquake loads. A commercial computer program called PLS-Tower [19] was used to create the tower geometry, to apply the loads on the tower, to analyze the tower under loads, and to specify the member sizes. In the prototype design, S355 grade steel ( Fy = 355 MPa) was used for the angle sections, and 20 mm diameter grade 8.8 bolts (Fy = 640 MPa, Fy = 800 MPa) were used for the connections. Each main leg in the prototype tower has two L150x16 angle sections positioned corner-to-corner below the investigated connection, while a single L160x17 angle section is used above this connection. Due to the limitations on the load capacity of the test setup, the test specimens used in the study were fabricated with a scale factor of 1:2.7, resulting in L60x6 angle sections and 8 mm diameter bolts. Testing of transmission tower subassemblies with scaled down models has previously been adopted by several researchers [8–10,20,21]. For the 8 mm diameter bolts, holes of 10 mm in diameter were used in members. The test specimens were fabricated by a local manufacturer of overhead transmission line tower supplier following the manufacturing methods conventionally used for the manufacturing of these towers. The connection detail that is currently used for the transition from a double-angle section to a single angle section is shown in Fig. 2, together with the other details considered in this study. The original
configuration (configuration A in Fig. 2) includes an interior and an exterior reinforcement angle, as well as a tie plate and two filler plates placed between the reinforcement angles and the main members. There are 16 bolts on each leg of the two lower main members, resulting in a total of 64 bolts in the connection. The two angle sections forming each main leg in the prototype tower are connected intermittently by 8 mm thick batten plates. Such an application results in an 8 mm gap between the angle sections of the main legs. To accommodate this gap, two 8 × 55 × 315 mm filler plates were used in the test specimens for configurations A–E. Each filler plate was replaced with two 8 mm thick plate washers for configuration F, as indicated in Fig. 2. The original connection detail also includes an 8 × 128 × 315 mm tie plate that is bolted to one leg of each of the two lower main member angles with 16 bolts. The tie plate was eliminated from the connection for configurations B, C, and F. Details of the investigated connection configurations are provided in Table 1 and the geometric details of angle members are given in Fig. 3. In Specimens A through E the bolts were tightened with an air impact wrench. The experience gained during fabrication of the test specimens indicates that the tightness of the bolts obtained with the air impact wrench approximately corresponded to that obtained by the full effort of a person using a hand wrench. Such application was adopted in order to reflect the bolt installation practice that is usually used in erection of steel lattice transmission towers. In Specimen F, on the other hand, the air impact wrench was used only on the four bolts that were located where plate washers were used. Because there is a gap of 8 mm between the legs of the upper and lower main member angles at other bolt locations, these bolts were left as “finger tight”.
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255
Fig. 2. Investigated connection configurations.
Tension tests were performed on coupon samples cut from the upper main member in Specimens A, C, D, and E. The measured yield and ultimate strength values from these tests are given in Table 2. 2.2. Test setup and procedure As shown in Fig.4, the lower end of the specimen was attached to the testing frame and vertical upward displacement was applied at the upper end with a hydraulic actuator. The top end of the upper main member angle was bolted to two pieces of steel plate that were welded to a horizontally positioned support plate. The support plate itself was attached to the swivel head of the hydraulic actuator. A similar detail was used at the lower end to fix the double-angle section of the lower main member to the testing frame. Displacement loading was applied with a constant rate of 0.02 mm/s until the failure of specimen occurred due to fracture of the upper main member. A displacement transducer was used to measure the deformation within the middle 50 cm length of the specimen. In addition to the displacement transducer, strain gages were used in some specimens on the main members and the interior reinforcement angle. The load cell at the end of the hydraulic cylinder, strain gages, and displacement transducer were connected to
Table 1 Details of connection configurations. Specimen
A
B
C
D
E
F
Interior reinf. angle Exterior reinf. angle Tie plate Filler plate
√ √ √ √
– √ √ √
√ – – √
– – – √
√ √
– – – –a
a
Each filler plate was replaced with two plate washers.
Fig. 3. Geometric details used in test specimens: (a) upper main member angle; (b) lower main member angle; (c) reinforcement angle.
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Table 2 Measured steel strength values. Specimen
A
C
D
E
Yield strength, MPa Ultimate strength, MPa
398 506
391 500
380 483
356 470
a data acquisition system and the readings were continuously collected and recorded during load tests. 2.3. Test results The failure of all specimens was due to complete net section fracture of the upper main member angle adjacent to the lead bolt holes (Fig. 5(a) and (b)). The fracture path went through four staggered bolt holes at both legs of the angle section. Inspection of the specimens indicated substantial necking and some level of bolt hole elongation at the fracture path. The upper main member also underwent significant amount of bending deformation. As evident in Fig. 5(c), the bending deformation localized around the fracture path. Some level of bending deformation was also observed in the lower main members (Fig. 5(d)). Even though the design of members in steel lattice towers is usually based on the assumption of pure axial loading, additional bending moments often develop in these members mainly due to connection eccentricity. Likewise, the eccentricity between the centroid of the upper main member angle and that of the built-up section of the lower main member in the test specimens led to development of bending moments in these members. Bending deformations observed in the upper and lower main member angles of the test specimens are the result of these bending moments that are induced by the connection eccentricity. The measured load–deformation response of specimens is given in Fig. 6. Superimposed on the plots are the calculated load capacities of the upper main member angle corresponding to the limit states of (1) yielding on net section (YON), (2) fracture on net section (FON),
and (3) yielding on gross section (YOG) [22]. For Specimens A, C, D, and E the measured yield and ultimate tensile strength values obtained from the tension coupon tests were used in these calculations. For Specimens B and F, on the other hand, no such tests were performed, and the material strengths were taken as the average of those of the other four specimens. The load capacity corresponding to the limit state of YOG was calculated as follows: P n ¼ F y ∙Ag
ð1Þ
where, Fy is the steel yield strength and Ag is the gross cross-sectional area of angle member. For the limit state of FON, the following relation was used to determine the load capacity: P n ¼ F u ∙Ae
ð2Þ
where, Fu is the ultimate strength and Ae is the effective net crosssectional area, which is defined as follows: Ae ¼ U∙An
ð3Þ
where, U is the shear lag factor and An is the net cross-sectional area. The net cross-sectional area is determined considering the material loss due to bolt holes, as well as the stagger of the holes. Based on this definition, the net width of the cross section that was used to calculate the net cross-sectional are of the angle section in test specimens was determined by subtracting the total hole diameter, which was taken as 2 mm larger than the bolt diameter, from the section width and adding for each gage space the quantity s2/4g, where s is the hole spacing (in the direction of loading) and g is the gage distance (in the direction transverse to the loading). The load capacity for the limit state of YON was calculated as follows: P n ¼ F y ∙Ae :
Fig. 4. Details of test setup.
ð4Þ
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Fig. 5. Typical deformation observed in specimens: (a), (b) fracture along critical path; (c) bending in upper main member angle; (d) bending in lower main member angles.
As evident in the plots given in Fig. 6, the calculated load capacity for the limit state of yielding on net section does not correspond to any specific physical event on the experimentally determined load–deformation curves. There is a good agreement, however, between the calculated net section fracture capacity and the maximum load attained by specimens during load tests. For Specimens A and B the calculated capacities overpredict the measured capacities by 3.8% and 0.6%, respectively. For the remaining four specimens, the calculated capacities are smaller than the measured capacities by a maximum of 6.5%. The experimentally determined maximum load capacity of specimens as well as the calculated capacities are summarized in Table 3. All connection configurations resulted in similar load–deformation response that includes an initial linear behavior followed by a yielding
region and a secondary linear portion that continues until the failure of specimens. The sudden increase of deformation observed in the response of some of the specimens is due to the relative slip occurring between the upper and lower main members within the connection region. The reason for having relative slip only in some specimens is most likely due to the difference among specimens in bolt pretension and the relative location of the bolts inside the bolt holes. During the fabrication of specimens, the bolts were tightened without any intent of applying pretension. Unintentional over-tightening of the bolts in some specimens might have resulted in higher resistance against relative slip in these specimens. Due to the fact that the holes in the upper and lower main member angles were 2 mm oversized, some level of relative slip would be expected to occur during load testing of
Fig. 6. Load–displacement response of specimens.
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Table 3 Measured and predicted load capacities. Specimen
A
B
C
D
E
F
Measured capacity, kN
254.4 207.8 264.1 275.0 291.6
257.6 201.0 259.3 266.0 275.5
270.3 204.1 261.0 270.2 288.9
269.6 198.4 252.1 262.6 268.0
253.2 185.8 245.3 246.0 257.6
272.5 201.0 259.3 266.0 269.3
Calculated capacity, kN
Y.O.N. F.O.N. Y.O.G.
Capacity from FEM, kN
specimens. However, the magnitude of such relative slip is dependent on the position of the bolts inside bolt holes. Therefore, any difference in the relative position of the bolts among the specimens would also result in different amounts of relative slip. A comparison of the load–deformation plots given in Fig. 6 also suggests that the occurrence of bolt slip does not have an unfavorable effect on the load capacity of specimens. Initial stiffness of specimens, as represented by the slope of the initial linear portion of load–deformation curves given in Fig. 6, is very similar. There are also minor differences among the slope of the post-yield portion of the curves and the maximum load capacities. It should be noted here that the connection details in some of the specimens were much more complicated than the others, with the total number of bolts being either 64, 48, or 32. As indicated by the results of the load tests, these major differences in the connection details did not cause any appreciable difference in the behavior of specimens. Based on this observation, it could be stated that the interior and exterior reinforcement angles as well as the tie plate used in the original connection detail could be eliminated and the total number of bolts could be reduced by a half without causing a major change in the behavior of the connection under tensile loading.
3. Numerical studies 3.1. Finite element models Numerical analysis of the investigated connections was conducted using the finite element (FE) models of test specimens prepared in the commercially available computer program ANSYS [23]. The connection geometries used in the finite element models were the same as those used in the experimental part of the study. The finite element model of Specimen A is shown in Fig. 7. Eight node solid elements with three displacement degrees of freedom at each node were used to model the angles, plates, and bolts in specimens. The elasto-plastic multilinear stress–strain relation shown in Fig. 8 was used together with the von Mises yield criterion to represent the material behavior in angles and plates. The modulus of elasticity and Poisson's ratio of steel were taken, respectively as 200 GPa and 0.3.
Fig. 7. Typical finite element model of test specimens.
Fig. 8. Engineering stress–strain relation used in models.
Additional effort was spent in order to accurately model the interaction between different components in models. Contact surfaces were defined at interfaces where there is force transfer between two different components. This includes the interface between (1) the angles and plates, (2) the bolt shank and the connecting elements, (3) the bolt heads and the connecting elements. These contact surfaces allow the transfer of friction and normal force between contact surfaces, while preventing the penetration of components into each other. The contact and target elements available in ANSYS were defined at the required locations to obtain proper contact behavior in the connection region. The value of friction and the level of bolt pretension are the two major factors that affect the slip behavior in the investigated connections. A coefficient of friction value of 0.35 was used in the analyses assuming Class A faying surfaces as described in the AISC 360-05 specification [22]. The initial pretension stress in the bolts was defined in the shank part of each bolt using a special element that is available in ANSYS for this purpose. As explained in the following part, two different levels of bolt pretension were used in the analyses. In order to cover several possible practices of tower installation, each model was analyzed with two cases: (1) slip case and (2) no slip case. In the slip case, 2 mm oversize bolt holes were used, i.e., the bolt hole diameter was taken as 10 mm. The orientations of the connection parts were adjusted such that each bolt was exactly centered inside the bolt hole. In this case, the bolts were assigned an initial pretension force of 10 kN. According to the AISC 360-05 specification the minimum bolt pretension is equal to 70% of the minimum tensile strength of bolts for full pretension connections [22]. In addition, a multiplier of 1.13 is defined to reflect the ratio of the mean installed bolt pretension to the specified minimum bolt pretension. Using these values results in a pretension force of approximately 32 kN for the bolts used in the connection parts of the test specimens. Considering this level of pretension together with the fact that the bolts in the test specimens were tightened with an air impact wrench, using 10 kN of bolt pretension in the finite element models was deemed suitable. In the no slip case, the relative slip between the connection components was prevented by using 10 mm diameter bolts inside 10 mm diameter holes. Because no relative slip was allowed, the value of initial bolt pretension was actually unimportant for the no slip case. For this reason, a nominal pretension of 1 kN was used in the bolts. The free ends of the two lower main member angles were fixed by restraining all three displacement components of the nodes located at the cross section. Loading was specified in the form of monotonically increasing displacement applied at the end of the upper main member angle. For this purpose, a rigid block was defined at the free end of the upper main member angle, and the displacement loading was defined at the center of this block.
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Fig. 9. Comparison of experimentally observed and predicted deformation patterns: (a) location of fracture plane in upper main member angle; (b) bending deformation of upper main member angle; (c) tilting of bolts in Specimen F.
3.2. Analysis results The FE models were able to capture the experimentally observed deformation patterns of test specimens. Formation of necking and net section fracture in the upper main member angle adjacent to the lead bolts as predicted by the FE models is shown in Fig. 9(a). The models were also able to predict the overall bending deformation of the upper main member angle (Fig. 9(b)). The replacement of filler plates with plate washers in Specimen F resulted in tilting of bolts prior to net section fracture of the upper main member angle during load tests. As shown in Fig. 9(c), a similar deformation mode is observed in the FE analysis of this specimen. Measured and predicted load–deformation responses of all connection configurations are presented in Fig. 10. For the predicted response, the results from FE models with slip and no slip cases as described earlier are given in the figure. For the friction coefficient and bolt pretension values used in the models, bolt slip occurred at around 200 kN of tensile load in connection configurations A through E. For connection configuration F, on the other hand, bolt slip occurred at a much smaller
load level because the filler plates were replaced by four pieces of plate washers in this configuration. Occurrence of bolt slip resulted in an increase of deformation while the load remains approximately constant. Following the closing of the gap between the bolt and side of the bolt hole, a contact condition is achieved and the load resisted by the connection starts to increase again. No such deformation is observed in the no slip case, as the bolt and hole diameters were the same in these models. Even though the occurrence of bolt slip in FE models resulted in load–deformation curves that look different than those for the no slip case, the two companion models of each connection configuration were able to reach the same level of maximum load capacity. A comparison of the measured and predicted curves presented in Fig. 10 suggest that the results from the model assuming bolt slip agree better with the experimentally obtained response than the results of the model assuming no slip, except for connection configuration F. Among all connection configurations studied, the largest difference between the experimentally determined and FEM predicted load capacities is 15% (for configuration A), while the smallest difference is
Fig. 10. Comparison of measured and predicted response.
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Fig. 11. Variation of strain on interior reinforcement angle (configuration A).
1% (for configurations D and F). Load capacities predicted by the FE model of each connection configuration are summarized in Table 3. The level of agreement between the response predicted by the FE model incorporating bolt slip and the experimentally determined response validates the combination of bolt pretension and slip coefficient values used in the analyses. Variation of strain on the interior reinforcement angle used in connection configuration A is shown in Fig. 11. The strain values in the test specimens were obtained from strain gages placed on the reinforcement angle at the two locations indicated in the figure. Positive and negative signs indicate tensile and compressive strain, respectively. As depicted on the plots, under the tensile load applied on the specimen, the strain on the interior reinforcement angle is mostly tensile at the midlength of the angle, while the strain at the upper end of the angle is always compressive. The compressive strains are the result of bending deformations that occurred in the specimens. As explained earlier, significant amount of bending deformation was observed on the upper main member angle and this deformation was mostly localized around the lead bolt holes. It should be noted that the strain gage that indicates compressive strain on the interior reinforcement angle was located adjacent to the lead bolt holes. Similar to the case with the overall load–deformation response, the variation of strains on the interior reinforcement angle is better predicted by the FE model that incorporates bolt slip as compared to the model that assumes no slip. 4. Conclusions Behavior of a bolted splice connection used in main legs of steel lattice transmission towers was studied experimentally and numerically. The experimental part of the study included tensile load testing of six specimens with different connection details. The main parameters used in the testing program were the number of bolts used in the connection as well as the presence of connection reinforcement angles and tie plate. For all connection configurations studied, the failure occurred due to net section fracture of the upper main member angle near the leading bolt holes. The calculated load capacity based on the measured material strength closely predicted the measured load capacity of specimens. Strain measurements and observed deformation shapes indicated the presence of bending in different members. Significant bending deformation was observed in the upper main member angle near the leading bolt holes. Results from 3-dimensional nonlinear FE analyses indicated similar response as the test specimens in terms of deformation patterns and load capacities. For each connection configuration studied, the experimentally determined response was better predicted by the FE model that incorporates bolt slip as compared to the model that assumes no slip.
The experimental and numerical results indicate that major differences among the investigated connection details do not cause any appreciable difference in the behavior under tensile loading. Based on this observation, it could be concluded that reinforcing angles and tie plate used in the original connection detail could be eliminated and the total number of bolts in the connection could be reduced by a half without causing a major change in the behavior of connection under tensile loading. References [1] N.P. Rao, G.M.S. Knight, N. Lakshmanan, Studies on failure of transmission line towers in testing, Eng. Struct. 35 (2012) 55–70. [2] F. Albermani, S. Kitipornchai, R.W.K. Chan, Failure analysis of transmission towers, Eng. Fail. Anal. 16 (2009) 1922–1928. [3] P.S. Lee, G. McClure, Elastoplastic large deformation analysis of a lattice steel tower structure and comparison with full-scale tests, J. Constr. Steel Res. 63 (2007) 709–717. [4] N.P. Rao, V. Kalyanaraman, Non-linear behaviour of lattice panel of angle towers, J. Constr. Steel Res. 57 (2001) 1337–1357. [5] N.P. Rao, G.M.S. Knight, N. Lakshmanan, N.R. Iyer, Effect of bolt slip on tower deformation, Pract. Period. Struct. Des. Constr. 17 (2012) 65–73. [6] S. Kitipornchai, F.G.A. Al-Bermani, A.H. Peyrot, Effect of bolt slippage on ultimate behavior of lattice structures, J. Struct. Eng. 120 (1994) 2281–2287. [7] W.Q. Jiang, Z.Q. Wang, G. McClure, G.L. Wang, J.D. Geng, Accurate modeling of joint effects in lattice transmission towers, Eng. Struct. 33 (2011) 1817–1827. [8] C. Lu, X. Ma, J.E. Mills, The structural effect of bolted splices on retrofitted transmission tower angle members, J. Constr. Steel Res. 95 (2014) 263–278. [9] Q. Xie, L. Sun, Experimental study on the mechanical behavior and failure mechanism of a lattice steel transmission tower, J. Struct. Eng. 139 (2013) 1009–1018. [10] F. Albermani, M. Mahendran, S. Kitipornchai, Upgrading of transmission towers using a diaphragm bracing system, Eng. Struct. 26 (2004) 735–744. [11] Y. Zhuge, J.E. Mills, X. Ma, Modelling of steel lattice tower angle legs reinforced for increased load capacity, Eng. Struct. 43 (2012) 160–168. [12] C. Lu, X. Ma, J.E. Mills, Modeling of retrofitted steel transmission towers, J. Constr. Steel Res. 112 (2015) 138–154. [13] T.D. Hoang, C. Herbelot, A. Imad, On failure mode analysis in a bolted single lap joint under tension-shearing, Eng. Fail. Anal. 24 (2012) 9–25. [14] P. Prabha, S.A. Jayachandran, M. Saravanan, V. Marimuthu, Prediction of the tensile capacity of cold formed angles experiencing shear lag, Thin-Walled Struct. 49 (2011) 1348–1358. [15] P. Moze, D. Beg, High strength steel tension splices with one or two bolts, J. Constr. Steel Res. 66 (2010) 1000–1010. [16] V.F. Paula, M.B. Luciano, W.T. Matias, Efficiency reduction due to shear lag on bolted cold-formed steel angles, J. Constr. Steel Res. 64 (2008) 571–583. [17] K.E. Barth, J.G. Orbison, R. Nukala, Behavior of steel tension members subjected to uniaxial loading, J. Constr. Steel Res. 58 (2002) 1103–1120. [18] R. Puthli, O. Fleischer, Investigations on bolted connections for high strength steel members, J. Constr. Steel Res. 57 (2001) 313–326. [19] Powerline Systems Inc., PLS-Tower Version 12.50 User's Manual, 2012. [20] J.E. Mills, X. Ma, Y. Zhuge, Experimental study on multi-panel retrofitted steel transmission towers, J. Constr. Steel Res. 78 (2012) 58–67. [21] D.E. O'Claire, D.M. Hesse, Transmission towers with cruciform legs, Electrical Transmission and Substation Structures 2012, pp. 275–287. [22] American Institute of Steel Construction, AISC 360-05 Specification for Structural Steel Buildings, AISC, 2005. [23] ANSYS, Version 15.0 Online User's Manual, 2013.
Contents lists available at ScienceDirect
Journal of Constructional Steel Research
Experimental and numerical analysis of a bolted connection in steel transmission towers Eray Baran a,⁎, Tolga Akis b, Gokmen Sen c, Ammar Draisawi b a b c
Department of Civil Engineering, Middle East Technical University, Ankara 06800, Turkey Department of Civil Engineering, Atilim University, Ankara 06836, Turkey MITENG, Ankara 06800, Turkey
a r t i c l e
i n f o
Article history: Received 9 July 2015 Received in revised form 21 January 2016 Accepted 12 February 2016 Available online xxxx Keywords: Lattice transmission tower Topic: Bolted connection Splice connection Steel angle Bolt slip Finite element analysis Tension member
a b s t r a c t This paper presents an integrated numerical and experimental study on a bolted splice connection used in main legs of steel lattice transmission towers. At specific locations, where the number of angle sections in built-up cross section of main leg members changes, the complex geometry around the connection region results in eccentricities in the load path and indirect load transfer. Such complex configurations and uncertainties in the load path have led to overdesigned connections with increased number of bolts and redundant connection reinforcing members. The current study was conducted in an attempt to gain a better understanding of the load-flow mechanism at this specific location where the cross section of main leg members changes. The experimental part included tensile load testing of six specimens with different connection details. The main parameters used in the testing program were the number of bolts used in the connection as well as the presence of connection reinforcement angles and tie plate. For all connection configurations studied, the failure occurred due to net section fracture of upper main member angle near leading bolt holes. The calculated load capacity based on the measured material strength closely predicted the measured load capacity of specimens. The experimentally determined response of each connection configuration was better predicted by the FE model that incorporates bolt slip as compared to the model that assumes no slip. The experimental and numerical results also indicate that major differences among the investigated connection details do not cause any appreciable difference in behavior under tensile loading. © 2016 Elsevier Ltd. All rights reserved.
1. Introduction Steel lattice towers have been extensively used worldwide as an important component of power transmission networks. These towers are usually constructed using equal leg angle section members and bearing type bolted connections. Depending on the magnitude of conductor and environmental loads as well as the geometric constraints of the transmission line, the axial force in main members at lower parts of the tower could get so large that built-up sections have to be used at these locations. These built-up sections are usually made of two, three, or four steel angle sections connected corner-to-corner, depending on the level of axial force in the member. At higher locations along the height of tower, axial force level in main legs decreases and this allows a reduction in the cross-sectional area of these members. Such reduction in the cross-sectional area of main legs is usually achieved by reducing the number of steel angle sections in the built-up cross section. At the specific locations, where the number of angle sections in the builtup cross section changes, the complex geometry around the connection region results in eccentricities in the load path and indirect load transfer ⁎ Corresponding author. E-mail address: [email protected] (E. Baran).
http://dx.doi.org/10.1016/j.jcsr.2016.02.009 0143-974X/© 2016 Elsevier Ltd. All rights reserved.
between several components. Such complex configurations and the uncertainties in the load path have led to overdesigned connections with increased number of bolts and redundant connection reinforcing members. The current study was conducted in an attempt to gain a better understanding of the load-flow mechanism at this specific location where the cross section of main leg members changes. The overall goal was to simplify the connection geometry by identifying the redundant members that can be eliminated from the connection without significantly affecting the structural behavior. An experimental study integrated with numerical analyses was conducted in order to achieve this goal. Numerical analysis of steel lattice towers has recently been performed by many researchers [1–4]. These numerical analyses were conducted using finite element models constructed using either one-dimensional beam–column elements or two-dimensional shell and plate elements. Eccentricity of connections and flexibility of joints were also considered in some of the studies. In majority of the cases, results obtained from the models were compared to the observed response from full-scale proof-load tests or smaller scale laboratory tests of the towers. These studies indicate that the maximum load capacity and the distribution of member failures can be predicted with a fair accuracy using numerical models. Based on the analysis of transmission
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line tower failures during prototype load testing, the possible reasons for tower failures are given as incorrect design assumptions, improper detailing, material defects, fabrication errors, force fitting during erection, and variation in bolt strength [1]. Several studies have also focused on the effect of bolt slip on nonlinear behavior of lattice towers [5–7]. Results from these studies indicated that the joint slip would be expected to affect the tower displacement, but have no major influence on the failure mode and ultimate strength of the tower. Another current area of research related with overhead transmission line towers is on strengthening of existing towers. Various strengthening methods to improve the load capacity of existing lattice towers so that they can tolerate the increase in conductor loads have recently been studied both experimentally and numerically [8–12]. These studies revealed that upgrading of existing towers could be achieved by providing additional bracing for members or by increasing the total cross-sectional area of individual members by reinforcement. Experimental and numerical studies on bolted splice connections of different types of metallic members are also available in the current literature [13–18].These studies revealed information regarding many aspects of bolted connections that can be useful in design of various bolted connections in steel lattice towers. The main phenomena investigated with these studies are the influence of connection geometry on the extent of shear lag effect on steel angles, the influence of the amount of edge distance as well as the level of bolt pretension on the behavior of plate connections, and the influence of connection eccentricity on the behavior of T-section tension members. 2. Experimental program The experimental part of the study included tensile load testing of six specimens with different connection details. The main parameters used in the testing program were the number of bolts used in the connection as well as the presence of connection reinforcement angles and tie plate. The specimens were fabricated at the production facility of a local manufacturer of transmission line towers, and were tested at Atilim University Structural Mechanics Laboratory. Fig. 1. Prototype design of 75 m tall 500 kV capacity tower.
2.1. Test specimens A prototype latticed transmission tower was designed, and the splice connection at main legs of this prototype structure was used as a basis for the test specimens. General geometry of the prototype tower and the location of the investigated splice connection along the length of the tower are shown in Fig. 1. The loads considered during design of the 75 m tall 500 kV prototype tower included wind, ice, equipment, conductor, broken wire, and earthquake loads. A commercial computer program called PLS-Tower [19] was used to create the tower geometry, to apply the loads on the tower, to analyze the tower under loads, and to specify the member sizes. In the prototype design, S355 grade steel ( Fy = 355 MPa) was used for the angle sections, and 20 mm diameter grade 8.8 bolts (Fy = 640 MPa, Fy = 800 MPa) were used for the connections. Each main leg in the prototype tower has two L150x16 angle sections positioned corner-to-corner below the investigated connection, while a single L160x17 angle section is used above this connection. Due to the limitations on the load capacity of the test setup, the test specimens used in the study were fabricated with a scale factor of 1:2.7, resulting in L60x6 angle sections and 8 mm diameter bolts. Testing of transmission tower subassemblies with scaled down models has previously been adopted by several researchers [8–10,20,21]. For the 8 mm diameter bolts, holes of 10 mm in diameter were used in members. The test specimens were fabricated by a local manufacturer of overhead transmission line tower supplier following the manufacturing methods conventionally used for the manufacturing of these towers. The connection detail that is currently used for the transition from a double-angle section to a single angle section is shown in Fig. 2, together with the other details considered in this study. The original
configuration (configuration A in Fig. 2) includes an interior and an exterior reinforcement angle, as well as a tie plate and two filler plates placed between the reinforcement angles and the main members. There are 16 bolts on each leg of the two lower main members, resulting in a total of 64 bolts in the connection. The two angle sections forming each main leg in the prototype tower are connected intermittently by 8 mm thick batten plates. Such an application results in an 8 mm gap between the angle sections of the main legs. To accommodate this gap, two 8 × 55 × 315 mm filler plates were used in the test specimens for configurations A–E. Each filler plate was replaced with two 8 mm thick plate washers for configuration F, as indicated in Fig. 2. The original connection detail also includes an 8 × 128 × 315 mm tie plate that is bolted to one leg of each of the two lower main member angles with 16 bolts. The tie plate was eliminated from the connection for configurations B, C, and F. Details of the investigated connection configurations are provided in Table 1 and the geometric details of angle members are given in Fig. 3. In Specimens A through E the bolts were tightened with an air impact wrench. The experience gained during fabrication of the test specimens indicates that the tightness of the bolts obtained with the air impact wrench approximately corresponded to that obtained by the full effort of a person using a hand wrench. Such application was adopted in order to reflect the bolt installation practice that is usually used in erection of steel lattice transmission towers. In Specimen F, on the other hand, the air impact wrench was used only on the four bolts that were located where plate washers were used. Because there is a gap of 8 mm between the legs of the upper and lower main member angles at other bolt locations, these bolts were left as “finger tight”.
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Fig. 2. Investigated connection configurations.
Tension tests were performed on coupon samples cut from the upper main member in Specimens A, C, D, and E. The measured yield and ultimate strength values from these tests are given in Table 2. 2.2. Test setup and procedure As shown in Fig.4, the lower end of the specimen was attached to the testing frame and vertical upward displacement was applied at the upper end with a hydraulic actuator. The top end of the upper main member angle was bolted to two pieces of steel plate that were welded to a horizontally positioned support plate. The support plate itself was attached to the swivel head of the hydraulic actuator. A similar detail was used at the lower end to fix the double-angle section of the lower main member to the testing frame. Displacement loading was applied with a constant rate of 0.02 mm/s until the failure of specimen occurred due to fracture of the upper main member. A displacement transducer was used to measure the deformation within the middle 50 cm length of the specimen. In addition to the displacement transducer, strain gages were used in some specimens on the main members and the interior reinforcement angle. The load cell at the end of the hydraulic cylinder, strain gages, and displacement transducer were connected to
Table 1 Details of connection configurations. Specimen
A
B
C
D
E
F
Interior reinf. angle Exterior reinf. angle Tie plate Filler plate
√ √ √ √
– √ √ √
√ – – √
– – – √
√ √
– – – –a
a
Each filler plate was replaced with two plate washers.
Fig. 3. Geometric details used in test specimens: (a) upper main member angle; (b) lower main member angle; (c) reinforcement angle.
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Table 2 Measured steel strength values. Specimen
A
C
D
E
Yield strength, MPa Ultimate strength, MPa
398 506
391 500
380 483
356 470
a data acquisition system and the readings were continuously collected and recorded during load tests. 2.3. Test results The failure of all specimens was due to complete net section fracture of the upper main member angle adjacent to the lead bolt holes (Fig. 5(a) and (b)). The fracture path went through four staggered bolt holes at both legs of the angle section. Inspection of the specimens indicated substantial necking and some level of bolt hole elongation at the fracture path. The upper main member also underwent significant amount of bending deformation. As evident in Fig. 5(c), the bending deformation localized around the fracture path. Some level of bending deformation was also observed in the lower main members (Fig. 5(d)). Even though the design of members in steel lattice towers is usually based on the assumption of pure axial loading, additional bending moments often develop in these members mainly due to connection eccentricity. Likewise, the eccentricity between the centroid of the upper main member angle and that of the built-up section of the lower main member in the test specimens led to development of bending moments in these members. Bending deformations observed in the upper and lower main member angles of the test specimens are the result of these bending moments that are induced by the connection eccentricity. The measured load–deformation response of specimens is given in Fig. 6. Superimposed on the plots are the calculated load capacities of the upper main member angle corresponding to the limit states of (1) yielding on net section (YON), (2) fracture on net section (FON),
and (3) yielding on gross section (YOG) [22]. For Specimens A, C, D, and E the measured yield and ultimate tensile strength values obtained from the tension coupon tests were used in these calculations. For Specimens B and F, on the other hand, no such tests were performed, and the material strengths were taken as the average of those of the other four specimens. The load capacity corresponding to the limit state of YOG was calculated as follows: P n ¼ F y ∙Ag
ð1Þ
where, Fy is the steel yield strength and Ag is the gross cross-sectional area of angle member. For the limit state of FON, the following relation was used to determine the load capacity: P n ¼ F u ∙Ae
ð2Þ
where, Fu is the ultimate strength and Ae is the effective net crosssectional area, which is defined as follows: Ae ¼ U∙An
ð3Þ
where, U is the shear lag factor and An is the net cross-sectional area. The net cross-sectional area is determined considering the material loss due to bolt holes, as well as the stagger of the holes. Based on this definition, the net width of the cross section that was used to calculate the net cross-sectional are of the angle section in test specimens was determined by subtracting the total hole diameter, which was taken as 2 mm larger than the bolt diameter, from the section width and adding for each gage space the quantity s2/4g, where s is the hole spacing (in the direction of loading) and g is the gage distance (in the direction transverse to the loading). The load capacity for the limit state of YON was calculated as follows: P n ¼ F y ∙Ae :
Fig. 4. Details of test setup.
ð4Þ
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Fig. 5. Typical deformation observed in specimens: (a), (b) fracture along critical path; (c) bending in upper main member angle; (d) bending in lower main member angles.
As evident in the plots given in Fig. 6, the calculated load capacity for the limit state of yielding on net section does not correspond to any specific physical event on the experimentally determined load–deformation curves. There is a good agreement, however, between the calculated net section fracture capacity and the maximum load attained by specimens during load tests. For Specimens A and B the calculated capacities overpredict the measured capacities by 3.8% and 0.6%, respectively. For the remaining four specimens, the calculated capacities are smaller than the measured capacities by a maximum of 6.5%. The experimentally determined maximum load capacity of specimens as well as the calculated capacities are summarized in Table 3. All connection configurations resulted in similar load–deformation response that includes an initial linear behavior followed by a yielding
region and a secondary linear portion that continues until the failure of specimens. The sudden increase of deformation observed in the response of some of the specimens is due to the relative slip occurring between the upper and lower main members within the connection region. The reason for having relative slip only in some specimens is most likely due to the difference among specimens in bolt pretension and the relative location of the bolts inside the bolt holes. During the fabrication of specimens, the bolts were tightened without any intent of applying pretension. Unintentional over-tightening of the bolts in some specimens might have resulted in higher resistance against relative slip in these specimens. Due to the fact that the holes in the upper and lower main member angles were 2 mm oversized, some level of relative slip would be expected to occur during load testing of
Fig. 6. Load–displacement response of specimens.
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Table 3 Measured and predicted load capacities. Specimen
A
B
C
D
E
F
Measured capacity, kN
254.4 207.8 264.1 275.0 291.6
257.6 201.0 259.3 266.0 275.5
270.3 204.1 261.0 270.2 288.9
269.6 198.4 252.1 262.6 268.0
253.2 185.8 245.3 246.0 257.6
272.5 201.0 259.3 266.0 269.3
Calculated capacity, kN
Y.O.N. F.O.N. Y.O.G.
Capacity from FEM, kN
specimens. However, the magnitude of such relative slip is dependent on the position of the bolts inside bolt holes. Therefore, any difference in the relative position of the bolts among the specimens would also result in different amounts of relative slip. A comparison of the load–deformation plots given in Fig. 6 also suggests that the occurrence of bolt slip does not have an unfavorable effect on the load capacity of specimens. Initial stiffness of specimens, as represented by the slope of the initial linear portion of load–deformation curves given in Fig. 6, is very similar. There are also minor differences among the slope of the post-yield portion of the curves and the maximum load capacities. It should be noted here that the connection details in some of the specimens were much more complicated than the others, with the total number of bolts being either 64, 48, or 32. As indicated by the results of the load tests, these major differences in the connection details did not cause any appreciable difference in the behavior of specimens. Based on this observation, it could be stated that the interior and exterior reinforcement angles as well as the tie plate used in the original connection detail could be eliminated and the total number of bolts could be reduced by a half without causing a major change in the behavior of the connection under tensile loading.
3. Numerical studies 3.1. Finite element models Numerical analysis of the investigated connections was conducted using the finite element (FE) models of test specimens prepared in the commercially available computer program ANSYS [23]. The connection geometries used in the finite element models were the same as those used in the experimental part of the study. The finite element model of Specimen A is shown in Fig. 7. Eight node solid elements with three displacement degrees of freedom at each node were used to model the angles, plates, and bolts in specimens. The elasto-plastic multilinear stress–strain relation shown in Fig. 8 was used together with the von Mises yield criterion to represent the material behavior in angles and plates. The modulus of elasticity and Poisson's ratio of steel were taken, respectively as 200 GPa and 0.3.
Fig. 7. Typical finite element model of test specimens.
Fig. 8. Engineering stress–strain relation used in models.
Additional effort was spent in order to accurately model the interaction between different components in models. Contact surfaces were defined at interfaces where there is force transfer between two different components. This includes the interface between (1) the angles and plates, (2) the bolt shank and the connecting elements, (3) the bolt heads and the connecting elements. These contact surfaces allow the transfer of friction and normal force between contact surfaces, while preventing the penetration of components into each other. The contact and target elements available in ANSYS were defined at the required locations to obtain proper contact behavior in the connection region. The value of friction and the level of bolt pretension are the two major factors that affect the slip behavior in the investigated connections. A coefficient of friction value of 0.35 was used in the analyses assuming Class A faying surfaces as described in the AISC 360-05 specification [22]. The initial pretension stress in the bolts was defined in the shank part of each bolt using a special element that is available in ANSYS for this purpose. As explained in the following part, two different levels of bolt pretension were used in the analyses. In order to cover several possible practices of tower installation, each model was analyzed with two cases: (1) slip case and (2) no slip case. In the slip case, 2 mm oversize bolt holes were used, i.e., the bolt hole diameter was taken as 10 mm. The orientations of the connection parts were adjusted such that each bolt was exactly centered inside the bolt hole. In this case, the bolts were assigned an initial pretension force of 10 kN. According to the AISC 360-05 specification the minimum bolt pretension is equal to 70% of the minimum tensile strength of bolts for full pretension connections [22]. In addition, a multiplier of 1.13 is defined to reflect the ratio of the mean installed bolt pretension to the specified minimum bolt pretension. Using these values results in a pretension force of approximately 32 kN for the bolts used in the connection parts of the test specimens. Considering this level of pretension together with the fact that the bolts in the test specimens were tightened with an air impact wrench, using 10 kN of bolt pretension in the finite element models was deemed suitable. In the no slip case, the relative slip between the connection components was prevented by using 10 mm diameter bolts inside 10 mm diameter holes. Because no relative slip was allowed, the value of initial bolt pretension was actually unimportant for the no slip case. For this reason, a nominal pretension of 1 kN was used in the bolts. The free ends of the two lower main member angles were fixed by restraining all three displacement components of the nodes located at the cross section. Loading was specified in the form of monotonically increasing displacement applied at the end of the upper main member angle. For this purpose, a rigid block was defined at the free end of the upper main member angle, and the displacement loading was defined at the center of this block.
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Fig. 9. Comparison of experimentally observed and predicted deformation patterns: (a) location of fracture plane in upper main member angle; (b) bending deformation of upper main member angle; (c) tilting of bolts in Specimen F.
3.2. Analysis results The FE models were able to capture the experimentally observed deformation patterns of test specimens. Formation of necking and net section fracture in the upper main member angle adjacent to the lead bolts as predicted by the FE models is shown in Fig. 9(a). The models were also able to predict the overall bending deformation of the upper main member angle (Fig. 9(b)). The replacement of filler plates with plate washers in Specimen F resulted in tilting of bolts prior to net section fracture of the upper main member angle during load tests. As shown in Fig. 9(c), a similar deformation mode is observed in the FE analysis of this specimen. Measured and predicted load–deformation responses of all connection configurations are presented in Fig. 10. For the predicted response, the results from FE models with slip and no slip cases as described earlier are given in the figure. For the friction coefficient and bolt pretension values used in the models, bolt slip occurred at around 200 kN of tensile load in connection configurations A through E. For connection configuration F, on the other hand, bolt slip occurred at a much smaller
load level because the filler plates were replaced by four pieces of plate washers in this configuration. Occurrence of bolt slip resulted in an increase of deformation while the load remains approximately constant. Following the closing of the gap between the bolt and side of the bolt hole, a contact condition is achieved and the load resisted by the connection starts to increase again. No such deformation is observed in the no slip case, as the bolt and hole diameters were the same in these models. Even though the occurrence of bolt slip in FE models resulted in load–deformation curves that look different than those for the no slip case, the two companion models of each connection configuration were able to reach the same level of maximum load capacity. A comparison of the measured and predicted curves presented in Fig. 10 suggest that the results from the model assuming bolt slip agree better with the experimentally obtained response than the results of the model assuming no slip, except for connection configuration F. Among all connection configurations studied, the largest difference between the experimentally determined and FEM predicted load capacities is 15% (for configuration A), while the smallest difference is
Fig. 10. Comparison of measured and predicted response.
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Fig. 11. Variation of strain on interior reinforcement angle (configuration A).
1% (for configurations D and F). Load capacities predicted by the FE model of each connection configuration are summarized in Table 3. The level of agreement between the response predicted by the FE model incorporating bolt slip and the experimentally determined response validates the combination of bolt pretension and slip coefficient values used in the analyses. Variation of strain on the interior reinforcement angle used in connection configuration A is shown in Fig. 11. The strain values in the test specimens were obtained from strain gages placed on the reinforcement angle at the two locations indicated in the figure. Positive and negative signs indicate tensile and compressive strain, respectively. As depicted on the plots, under the tensile load applied on the specimen, the strain on the interior reinforcement angle is mostly tensile at the midlength of the angle, while the strain at the upper end of the angle is always compressive. The compressive strains are the result of bending deformations that occurred in the specimens. As explained earlier, significant amount of bending deformation was observed on the upper main member angle and this deformation was mostly localized around the lead bolt holes. It should be noted that the strain gage that indicates compressive strain on the interior reinforcement angle was located adjacent to the lead bolt holes. Similar to the case with the overall load–deformation response, the variation of strains on the interior reinforcement angle is better predicted by the FE model that incorporates bolt slip as compared to the model that assumes no slip. 4. Conclusions Behavior of a bolted splice connection used in main legs of steel lattice transmission towers was studied experimentally and numerically. The experimental part of the study included tensile load testing of six specimens with different connection details. The main parameters used in the testing program were the number of bolts used in the connection as well as the presence of connection reinforcement angles and tie plate. For all connection configurations studied, the failure occurred due to net section fracture of the upper main member angle near the leading bolt holes. The calculated load capacity based on the measured material strength closely predicted the measured load capacity of specimens. Strain measurements and observed deformation shapes indicated the presence of bending in different members. Significant bending deformation was observed in the upper main member angle near the leading bolt holes. Results from 3-dimensional nonlinear FE analyses indicated similar response as the test specimens in terms of deformation patterns and load capacities. For each connection configuration studied, the experimentally determined response was better predicted by the FE model that incorporates bolt slip as compared to the model that assumes no slip.
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