STRENGTH AND DEFORMATION CHARACTERISTICS OF TYPICAL X-LAM CONNECTIONS

Transcription

1 STRENGTH AND DEFORMATION CHARACTERISTICS OF TYPICAL X-LAM CONNECTIONS Igor Gavric 1, Massimo Fragiacomo and Ario Ceccotti 3 ABSTRACT: This paper presents some of the results of an extensive experimental programme on typical X-Lam connections, conducted at CNR-IVALSA research institute. The goal of this research is to provide a better understanding of the seismic performance of connections in cross-laminated timber buildings subjected to seismic actions. In-plane monotonic and cyclic shear tests were performed on mechanical screwed connections between adjacent parallel wall-wall and floor-floor X-Lam panels. In addition, monotonic and cyclic tests were carried out on orthogonally connected panels (wall-wall and wall-floor) subjected to shear and withdrawal load. Mechanical properties in terms of strength, stiffness, energy dissipation, ductility ratio and impairment of strength were evaluated. The overstrength factor, which is of great importance in capacity-based design, was also evaluated for the different types of connection tested. KEYWORDS: X-Lam panels, Cyclic tests, Joints, Mechanical fasteners, Strength and deformation characteristics 1 INTRODUCTION 13 The SOFIE research project ( Fiemme Construction System), coordinated and conducted by the CNR- IVALSA Trees and Timber Institute, started in 005 with an aim to develop seismic resistant sustainable multi-storey timber buildings made of prefabricated cross-laminated panels. First, racking tests of wall panels with different layouts of connections and openings [1] and pseudo-dynamic tests of a full scale one-storey building were performed [], continued with shaking table tests of a 3-storey building in 006 [3], and ended with shaking table tests of a 7-storey building in 007, the latter one conducted at the world s largest seismic testing facility (E-Defense) in Miki, Japan. Experimental tests provided excellent outcomes, as the buildings were able to survive a series of strong recorded earthquakes, such as Kobe earthquake (1995), virtually undamaged, while at the same time demonstrating significant energy dissipation. However, further research in this field is still needed in order to better characterize the seismic 1 Igor Gavric, PhD Candidate, DICAR Department of Civil Engineering and Architecture, University of Trieste. Piazzale Europa 1, 3417 Trieste, and Research Assistant, IVALSA Trees and Timber Institute, CNR. Via Biasi, San Michele all'adige (Tn). gavric.igor@gmail.com Massimo Fragiacomo, Associate Professor of Structural Design, Department of Architecture, Design and Urban Planning, University of Sassari. Piazza Duomo 6, Alghero. fragiacomo@uniss.it 3 Ario Ceccotti, Director, IVALSA Trees and Timber Institute, CNR. Via Biasi, San Michele all'adige (Tn). ceccotti@ivalsa.cnr.it behaviour of typical X-Lam connections (1-D models), single wall panels or series of adjacent wall panels (-D models), and entire X-Lam buildings (3-D models). In order to understand better the seismic performance of typical connections used in cross-lam buildings and their behaviour in lateral resistant X-Lam wall systems, an extended experimental programme on typical X-Lam connections was undertaken. Experimental values will be used to develop analytical predictions of the connection seismic behaviour in terms of strength and stiffness properties. Furthermore, experimental test results will serve for fitting newly developed hysteretic theoretical models [4]. The X-Lam connections will be represented as non-linear springs with hysteretic properties and will be modelled in advanced numerical software (Abaqus). Once non-linear components, representing X-Lam connections, are calibrated on the experimental results, the FE (-D) models will be validated on the experimental results of different wall configurations [1,5]. Furthermore, 3-D FE models of the single storey SOFIE building [], 3- and 7-storey SOFIE buildings will be implemented and used to compare numerical and experimental results. A parametric study will then be carried out using the aforementioned FE model to extend the results of the experimental tests to different configuration of technical interest, including different geometry of the buildings, types of accelerograms, values of PGA (peak ground acceleration), types of connectors, layout of connections, etc. In CLT construction, all walls in one storey contribute to the lateral and gravity resistance. Therefore, they provide a degree of redundancy and a system effect, which can

2 influence significantly the building stiffness and strength properties. The effect of wall-wall connections, wallfloor connections and floor-floor connections on the seismic performance of CLT walls is particularly influential as the X-Lam panels are relatively stiff, and therefore will be investigated in this paper. In a recent study [6,7] it was demonstrated that different approaches to model connections between panels in X-Lam buildings can markedly influence the building stiffness, vibration period, base shear forces, and building ductility. Thus, understanding the cyclic behaviour of connections in X-Lam building is crucial for their seismic design. Another important aspect of seismic design is to ensure the development of cyclic yielding occurs in the dissipative zones. All other structural members and connections shall be designed with sufficient overstrength. Dissipative zones shall be located in joints and connections, whereas the timber members themselves shall be regarded as behaving elastically [8]. Thus, overstrength values of typical X-Lam connection will be derived on the basis of statistical analysis of experimental test results on several samples of each configuration tested in the experimental programme. PREVIOUS RESEARCH ON X-LAM CONNECTIONS The research presented in this paper is part of an experimental programme on connections conducted at CNR IVALSA. The experimental programme is divided into three main parts: (i) hold-down connection tests; (ii) angle brackets connection tests; and (iii) screwed panelpanel connection tests. Only the latter tests are presented in this paper. Both hold-down and steel angle bracket tests were performed under monotonic and cyclic loading, separately in two directions: shear and tension [5]. Two different variations of connections were tested, steel base-panel and panel-panel connections. The first variation represents connections of wall panels to foundations connections. The second variation represents connections of wall panels to floor panels in upper stories of X-Lam buildings. All mechanical properties were derived following the pren151 standard [9] and the overstrength factors were evaluated using statistical analysis. For angle brackets the overstrength values in both directions, tension and shear, were found to range from 1.1 to 1.3, depending on the statistical approach. Values for hold-downs ranged from 1. to 1.3 in tension and from 1.4 to 1.6 in shear direction. An ongoing research project at Technical University of Graz (Austria) included a series of monotonic and cyclic tests on different types of angle brackets with several different types of anchoring. For each bracket at least three cyclic test repetitions were performed. Overstrength values were found to be below the value of 1.3 for both directions, shear and uplift [10]. Previously, a comprehensive investigation was undertaken at University of Ljubljana (Slovenia) to determine the seismic behaviour of X-Lam wall panels and connections [11]. In a preliminary study, an overstrength factor of 1.3 was suggested for angle brackets and 1.6 for screwed connections between cross-lam panels [7]. Experimental tests on several different types of screwed adjacent parallel wall panels were performed by Muñoz et. al. [1], Sandhaas et. al. [13] and Joyce et. al. [14]. Experimental values of load-carrying capacity were compared with calculations using existing design codes provisions (EC5 [0], CSA086 [1]) and with new design models proposed by Uibel and Blaß [15,16]. Their study focused on the development of a calculation methodology of the load-carrying capacity of connections with dowel-type fasteners in the direction to the CLT panel and parallel to their narrow side (i.e., edge joints). A simplified calculation model, minimum edge and end distances, and spacing of fasteners in the aforementioned connection types were proposed. In addition, withdrawal strength of selftapping screws, typically used in connecting X-Lam panels to the plane of the panel or in edge joints, was also investigated [17]. 3 EXPERIMENTAL TEST PROGRAMME For each of the 0 different configurations listed in Table 1 at least one monotonic and six cyclic tests were performed in order to obtain statistically representative values. The types of connections and type of X-Lam panels were selected based on the typical connection details used in the 3-story SOFIE building tested in Japan [3]. Table 1: Test matrix for para. wall-wall connection details Test configuration Loading direction Wall Wall panel connections (parallel panels) parallel (lap joint) parallel (spline joint) (lap joint) (spline joint) Type and number of screws HBS Φ8x80mm HBS Φ8x80mm x4 screws HBS Φ8x80 screws HBS Φ8x80

3 Table : Test matrix for perp. wall-wall connection details Test configuration Loading direction Wall Wall panel connections (orthogonal panels) Type and number of screws HBS Φ10x180 spacing between fasteners was in accordance with the reference building [18]. Table 4: Test matrix for floor-floor connection details Test configuration Floor Floor panel connections Loading direction Type and number of screws 13 HBS Φ10x parallel (lap joint) HBS Φ10x Withdrawal HBS Φ10x180 4 screws Screwed in-plane shear tests were performed on connections between parallel adjacent X-Lam panels, using different types of vertical joints: (i) spline joints with Kerto LVL strip, and (ii) lap joints. In addition, experimental tests were carried out also on orthogonally connected panels (wall-wall and wall-floor) subjected to shear in two different directions and withdrawal load (see Tables and 3). Table 3: Test matrix for wall-floor connection details Test configuration Wall Floor panel connections Loading direction Type and number of screws 0 4 TEST SETUPS (lap joint) HBS Φ10x140 screws Eurocode standard EN [8] states that the properties of dissipative zones of timber structures should be determined by tests performed either on single joints, on whole structures or on parts thereof in accordance with pren151 [9]. The standard procedure for cyclic testing of joints made with mechanical fasteners prescribed by pren151 was followed in all of the tests, with input displacement rate varying from 0. to 0.8 mm/s so that a duration of each test did not exceed the time limit of 30 min. Monotonic tests were carried out by displacement controlled ramping at a loading rate varying from 0.05 mm/s to 0. mm/s. Specimens were stored and tested under controlled conditions with 50% RH and 0 C. HBS Φ10x60 TEST 9: SECTION 1-1: HBS 8x80 FRONT VIEW: Withdrawal HBS Φ10x60 HBS Φ10x60 4 screws Finally, tests were performed on floor-floor panel connections, using lap joint (Table 4). In this paper 1 configurations of screwed panel-panel connection tests are presented (configurations 9-0). Wall panels were 5-layered X-Lam with a thickness of 85 mm ( ) and floor panels were 14 mm thick with 6-layered build-up structure ( ). Number of fasteners in each configuration and 85 TEST 10: TEST 19: HEA 0 profile fi16 threaded rods HBS 8x80 HBS 10x Figure 1: Test setups of configurations 9, 10 and 19 (measures in mm) All shear tests were conducted using a reversed cyclic procedure with predefined yield values which were varying from configuration to configuration, depending on experimental yield values obtained from monotonic tests. However, all tension tests and withdrawal tests were subjected to a non-reversed modification of the procedure outlined in pren151 due to restrained movement (timber-timber or timber-foundation contact) in compression direction. Shear tests of in-plane screwed connections between parallel adjacent panels were tested

4 with two different types: (i) lap joint with 50 mm overlap length for test configuration 9 and 10 mm long step joint for test configuration 19; and (ii) spline joint with Kerto LVL strip, notched into outer side of connection, 8 mm thick and 180 mm wide. In test configurations 9 and 19 panels were directly connected to each other with two self-tapping screws on each side, with 10 cm spacing between them. On the other hand, in case of configuration 10, double row of screws were used to connect the LVL strip with two wall panels, also with 10 cm spacing in vertical direction (Figure 1). Similarly, test configuration 11, 1 and 0 had screws positioned at a distance of 10 cm. While the lower panel was fixed to a HEA profile, the upper panel followed the non-reversed cyclic loading protocol. Over-lap dimensions and screw type in test 11 were equal as in configuration 9; test 1 was similar to test 10; and test 0 had the same connection detail as test 19 (Figure ). middle wall panel in the case of test 13 is rotated by 90 with respect to the panel in test 16. Consequently, the threaded parts of the screws in test 13 are positioned to the wood grain direction and in test 16 parallel to the wood grain direction. The layout of test 14 is similar to configuration 13 in terms of screw type, spacing, type of panels and orientation of panels, whereas configuration 17 is similar to configuration 16. The direction of loading in these two cases is at the shorter side of the panel edge. As the edge distances are relatively short, brittle failures of panels due to splitting might occur. In Figure 4 the orientation of the middle layer of wall panel with the threaded part of self-tapping screws can be seen. TEST 14: TEST 17: fi16 threaded rods TEST 11: TEST 1: SECTION 1-1: FRONT VIEW: TEST 0: HEA 0 profile fi16 threaded 1 rods HBS 10x180 HBS 10x HBS 8x HBS 10x Figure : Test setups of configurations 11, 1 and 0 (measures in mm) To investigate the cyclic shear behaviour of orthogonal wall-wall connections and wall-floor connections, two different types of test setups were designed. Whilst in all aforementioned test configurations the fasteners were positioned to the plane of the panels, in the following configurations (Figures 3-5) the fasteners were inserted into the narrow sides of the X-Lam panels, also known as edge joints TEST 13: TEST 16: HEA 0 profile fi16 threaded rods HBS 10x60 HBS 10x Figure 3: Test setups of configurations 13 and 16 (measures in mm) s 13 and 16 represent edge joints between wall-wall and wall-floor panels loaded in the direction of the longer side of the edge. On each side two self-tapping screws were positioned with 10 cm of spacing. As seen from Figure 3, the orientation of the Figure 4: Test setups of configurations 14 and 17 (measures in mm) To determine the withdrawal properties of X-Lam connections with self-tapping screws subjected to monotonic and cyclic loads, test configurations 15 and 18 were designed. In both cases four screws with 7.5 cm were used. In test configuration 15, screws were driven to the grain, while in case of configuration 18 screws were placed parallel to the grain TEST 15: TEST 18: fi16 threaded rods HBS 10x180 HBS 10x Figure 5: Test setups of configurations 15 and 18 (measures in mm) 5 RESULTS AND DISCUSSION The experimental test results were assessed in terms of strength, stiffness, energy dissipation, damping ratio and ductility, following the standard procedure from pren151 [9]

5 5.1 PARALLEL PANELS In-plane shear loading The hysteresis loops of configuration 19 obtained during the cyclic tests are displayed in Fig. 6 together with the 1 st, nd and 3 rd cycle backbone curves. Figure 7 shows a photo of the experimental setup of test configuration 19. Failure type of step joint connections and spline connections, subjected to in-plane shear was found to be mode e from European yielding model [0] with bending of the screws and formation of one plastic hinge. In some tests fracture of screws eventually occurred. A slight embedding of the screws head was noticed at higher rates of displacement. No brittle failure modes could be observed no cracks or splitting occurred during the monotonic and cyclic tests. Force [kn] 19 CS 04: Force Displacement Figure 6: Hysteresis loops of test configuration 19 Figure 7: Experimental setup of test configuration Displacement [mm] 1st cycle backbone curve nd cycle backbone curve 3rd cycle backbone curve Hysteresis Monotonic test Table 5 displays the average strength and deformation properties of connection configurations 9, 10 and 19, according to the pren151 standard. More specifically, k el and k pl represent initial stiffness and plastic stiffness; F y and v y signify yielding load and yielding displacement; F max and v max denote maximum load and maximum displacement; F u and v u signify ultimate load and ultimate displacement; D signifies ductility ratio (ratio between ultimate displacement and yield displacement); D mon represents ductility ratio, obtained from monotonic tests. As one of the criteria for ultimate value in pren151 is also strength at 30 mm displacement, force F 30 and ductility ratio D 30 are presented. Energy dissipation properties are measured by the quantities ν eq(1st) and ν eq(3rd), which represent the equivalent viscous damping ratio calculated at the 1 st and 3 rd cycles of the hysteretic loop at the displacement point where the maximum load was reached. In order to ensure that the given values of the behaviour factor may be used, ductility classes are defined in EC8-1. The dissipative zones shall be able to deform plastically for at least three fully reversed cycles at a static ductility ratio of 4 for ductility class M structures and at a static ductility ratio of 6 for ductility class H structures, without more than a 0% reduction of their resistance. The values of mechanical properties for a single specimen were analyzed by taking into account the results from both sides of hysteretic loops. The mean values for a specific connection configuration were obtained by calculating the average value of all six cyclic test results within the specific configuration. Based on a statistical analysis of the experimental results, the coefficients of variation (COV) were calculated (see Table 5). Table 5: Mechanical properties of parallel panel-panel connections under in-plane shear loading Mechanical property x mean COV x mean COV x mean COV k,el [kn/mm] k,pl [kn/mm] F y [kn] v y [mm] F max [kn] v max [mm] F u [kn] v u [mm] D [-] D mon [-] F 30 [kn] D 30 [-] Ductility class H H M ν eq(1st) [%] ν eq(3rd) [%] Step joint in wall-wall connection (Test 9) exhibited 50% higher initial stiffness in comparison with spline joint (Test 10). However, 40% higher resistance of test configuration 10 in comparison with test 9 was found at 46% higher maximum displacement. Furthermore, ultimate displacement at spline joint was 19% higher. Both test configurations 9 and 10 can be classified in ductility class H (high). In terms of energy dissipation capacity, spline joint exhibited 4% higher damping ratio at the 3 rd cycles. In test 19, a thicker panel, longer screws with larger diameter, and an over-lap with longer notch were used and resulted in higher yield load and yield displacement,

6 higher average strength but lower ductility level than in test configuration 9. In addition, damping ratio at the 3 rd cycles was found to be 66% higher In-plane axial loading The hysteresis loops of configuration 1 obtained during the cyclic tests are displayed in Fig. 8 together with the 1 st, nd and 3 rd cycle backbone curves. Figure 9 shows a photo of the experimental setup of test configuration 1. Failure modes associated with lap connections were found to be different than those with a LVL spline, when tested under in-plane axial loading. For connections with lap configuration (Tests 11 and 0), the mean failure mode was due to splitting of the wall panel or failure of glue bond in inner layers. In addition, in some tests plug shear in the zone of screw thread was observed. On the other hand, in cases of LVL spline connections the observed failure modes were mainly associated with pull-through of head with formation of one plastic hinge within the X-Lam panel (Test 1). Force [kn] CS 03: Force Displacement Displacement [mm] 1st cycle backbone curve nd cycle backbone curve 3rd cycle backbone curve Hysteresis Monotonic test Figure 8: Hysteresis loops of test configuration 1 Figure 9: Experimental setup of test configuration 1 By comparing the results from test configurations 11 and 1, it is possible to recognize a significant difference in initial stiffness. Step joint connection exhibits 33% higher initial stiffness in comparison with spline LVL connection. However, spline joint was able to resist higher forces at displacement values which were more than twice as large. The reason for that was mainly the different type of failure, as brittle failure modes in step joints prevented resistance to higher loads at larger displacements. In addition, COV (coefficient of variation) values of mechanical properties raised to 30%. 0 with 10 mm lap and 14 mm thick panel performed better than configuration 10 (50 mm lap, 85 mm thick panel) as no brittle failure modes were observed. Ultimate values were reached at relatively large displacements, on average more than 40 mm. Due to higher yielding displacement, the ductility ratio was relatively low and consequently so was also the ductility class. Energy dissipation capacity in the 1 st cycles was found to be noticeably higher, while nd and 3 rd cycles showed almost equal damping capacity as configuration 11. Table 6: Mechanical properties of parallel panel-panel connections under in-plane axial loading Mechanical property x mean COV x mean COV x mean COV k,el [kn/mm] k,pl [kn/mm] F y [kn] v y [mm] F max [kn] v max [mm] F u [kn] v u [mm] D [-] D mon [-] F 30 [kn] D 30 [-] Ductility class H H M ν eq(1st) [%] ν eq(3rd) [%] ORTHOGONAL PANELS 5..1 Shear loading parallel to the joint The hysteresis loops of configuration 16 obtained during the cyclic tests are displayed in Fig. 10 together with the 1 st, nd and 3 rd cycle backbone curves. Figure 11 shows a photo of the experimental setup of test configuration 16. Considerable embedment of the self-tapping screws head into the X-Lam panel was observed at higher displacement values. This could be an indication that the threaded length of the shank was able to hold the axial tensile forces generated in the screw without significant withdrawal, known as the rope effect which enhances the lateral load resistance of the connection. Displacements above 30 mm eventually caused splitting in the panel. As it can be seen in Table 7, mechanical properties for both configurations (13 and 16) were very similar to each other. This indicates that the orientation of the middle wall panel did not influence the overall performance of the aforementioned connection types.

7 Force [kn] 16 CS 07: Force Displacement Shear loading to the joint The hysteresis loops of configuration 17 obtained during the cyclic tests are displayed in Fig. 1 together with the 1 st, nd and 3 rd cycle backbone curves. Figure 13 shows a photo of the experimental setup of test configuration 17. Behaviour of test configurations 14 and 17 were similar in terms of mechanical response to cyclic loading (see Table 8) Displacement [mm] 1st cycle backbone curve nd cycle backbone curve 3rd cycle backbone curve Hysteresis Monotonic test 17 CS 07: Force Displacement Figure 10: Hysteresis loops of test configuration Force [kn] st cycle backbone curve nd cycle backbone curve 3rd cycle backbone curve Hysteresis Monotonic test -1 Displacement [mm] Figure 1: Hysteresis loops of test configuration 17 Figure 11: Experimental setup of test configuration 16 Table 7: Mechanical properties of orthogonal panel-panel connections under loading parallel to the joint Mechanical property x mean COV x mean COV k,el [kn/mm] k,pl [kn/mm] F y [kn] v y [mm] F max [kn] v max [mm] F u [kn] v u [mm] D [-] D mon [-] F 30 [kn] D 30 [-] Ductility class M M ν eq(1st) [%] ν eq(3rd) [%] Figure 13: Experimental setup of test configuration 17 One notable difference is the higher ultimate displacement reached in the case of test 17. It was found out that the reason for that was the failure modes. In both cases yielding of the screws occurred, leading to the splitting of the middle panel due to unfulfilled requirement for minimum thickness of X-Lam panel, which should have been at least 10 diameters according to the Uibel and Blaß proposal [16]. In the case of test 14, splitting occurred at lower values of displacement, leading to lower ultimate displacement attained.

8 Table 8: Mechanical properties of orthogonal panel-panel connections under loading to the joints Mechanical property x mean COV x mean COV k,el [kn/mm] k,pl [kn/mm] F y [kn] v y [mm] F max [kn] v max [mm] F u [kn] v u [mm] D [-] D mon [-] F 30 [kn] D 30 [-] Ductility class H H ν eq(1st) [%] ν eq(3rd) [%] Axial (withdrawal) loading The hysteresis loops of configuration 18 obtained during the cyclic tests are displayed in Fig. 14 together with the 1 st, nd and 3 rd cycle backbone curves. Figure 15 shows a photo of the experimental setup of test configuration 15. The modes of failure were mainly head penetration of screw head in the member representing the orthogonal wall element (Test 15) or floor element (Test 18). Sometimes it was combined with withdrawal from the wall panel. While turning back to the zero position after head embedment, local splitting occurred in the outer laminates of orthogonal panel due to forces acting to the grains. Force [kn] CS 04: Force Displacement Displacement [mm] 1st cycle backbone curve nd cycle backbone curve 3rd cycle backbone curve Hysteresis Monotonic test Figure 14: Hysteresis loops of test configuration 18 Figure 15: Experimental setup of test configuration 15 Both test configuration 15 and 18 performed with relatively low average yielding displacement, 1.30 mm for test 18 and 1. mm for test 15. Average yield force, peak force and ultimate force were all very similar in either case (Table 9), with difference of less than 10%. A reason for that could be the same type of failure mechanism. Penetration length was in both cases sufficient to resist the withdrawal forces. In most cases, penetration of the screw head led to pull-through the orthogonal panel. High level of ductility was therefore attained as in the post-elastic phase degradation of forces was relatively low. As expected, damping capacity due to screw withdrawal was relatively low, around 1.3-.% in the 3 rd cycles. Table 9: Mechanical properties of orthogonal panel-panel connections under withdrawal loading Mechanical property x mean COV x mean COV k,el [kn/mm] k,pl [kn/mm] F y [kn] v y [mm] F max [kn] v max [mm] F u [kn] v u [mm] D [-] D mon [-] F 30 [kn] D 30 [-] Ductility class H H ν eq(1st) [%] ν eq(3rd) [%]

9 5.3 OVERSTRENGTH FACTORS Brittle members in timber structures must be designed for the overstrength related to the strength of the ductile connections to ensure the ductile failure mechanism will take place before the failure of the brittle members. The ovestrength ratio γ ov is defined as the ratio between the 95 th percentile of the connection strength distribution and the analytical prediction of the design connection strength F d.. The design strength capacity F d was calculated by dividing the characteristic experimental strength F 0.05 by the strength partial factor γ M, assumed to be equal to one according to the Eurocode 8 for dissipative timber structures. The experimental characteristic strength values from tests were based on the lower 5 th percentile values assuming three different distributions, with a % confidence level; (i) normal distribution; (ii) log-normal distribution; (iii) standard procedure for calculation of characteristic 5-percentile values from the EN14358 standard [19]. A comparison of overstrength factors evaluation with these three approaches was done. With normal and lognormal distributions overstrength values were ranging from 1.15 to 1.7, except for the case of configuration 11, where due to the brittle failure, the high scatter of loadcarrying capacity leads to an overstrength factor of.3. An average overstrength value calculated including all 1 configurations was On the other hand, the approach from EN14358 standard gave the most conservative values, as this standard procedure requires larger variance for lower number of performed tests. Overstrength values range from 1. to 1.9, excluding the case of configuration 11, where the value was found to be 3.3. An average overstrength value of all test configurations using this procedure was Table 10: Overstrength factors (γ ov) of typical screwed X-Lam panel-panel connections Type of distribution Normal F 0.05 [kn] F 0.95 [kn] γ ov [-] Log-Normal F 0.05 [kn] F 0.95 [kn] γ ov [-] EN14358 F 0.05 [kn] F 0.95 [kn] γ ov [-] SUMMARY AND CONCLUSIONS An experimental testing programme on screwed connections between adjacent X-Lam panels was performed with the aim to better understand their performance when subjected to seismic actions. All mechanical properties in terms of pren151 procedure were evaluated. Step joint connections performed with higher stiffness in comparison with spline LVL joints but in some cases their failure modes was brittle due to splitting of the inner layer of the panel or plug shear. Spline joints could also resist greater forces at higher displacement values. The orientation of the layers in shear resistance of orthogonally connected panels through edge joint did not influence much the overall performance. Sufficient spacings, end distances, edge distances and panel thickness are required to prevent brittle failures. Withdrawal tests showed that the screw head penetration was more critical than the withdrawal of the screw from the panel. The outcomes of this connection test experimental programme will be used to calibrate advanced component FE models for non-linear static and dynamic numerical analyses of X-Lam walls and buildings, as well as to calibrate analytical calculation methods of X- Lam connections behavior. This research provides a part of the missing information on the overstrength factor for timber connections in the Eurocode 8 Timber part. ACKNOWLEDGEMENTS The research presented in this paper was funded by the Autonomous Province of Trento (Italy) within the SOFIE research project. The authors would like to acknowledge the contribution to this research of IVALSA laboratory staff Mario Pinna, Diego Magnago and Paolo Dellantonio, who provided technical expertise for the experimental testing. REFERENCES [1] Ceccotti A., Lauriola M., Pinna M., Sandhaas C. SOFIE Project - Cyclic Tests on Cross-Laminated Wooden Panels, Proceedings of the 9th World conference on timber engineering, Portland, Oregon (USA), August 6th-10th 006.

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