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Conventional and novel timber steel hybrid connections: testing, performance and assessment Schneider, Johannes 2015

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CONVENTIONAL AND NOVEL TIMBER STEEL HYBRID CONNECTIONS: TESTING, PERFORMANCE AND ASSESSMENT  by  Johannes Schneider  Dipl.-Ing. Bau, Universität Stuttgart, 2009   A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF  DOCTOR OF PHILOSOPHY  in  THE COLLEGE OF GRADUATE STUDIES  (Civil Engineering)  THE UNIVERSITY OF BRITISH COLUMBIA (Okanagan)     July 2015  © Johannes Schneider, 2015  ii Abstract  The combination of timber and steel for residential hybrid buildings is an unconventional approach to date. The challenge of a hybrid structure is the link between the individual materials. The focus of this research is the connection between steel frame and the infill wall. Over 100 conventional bracket-type connections with various combinations of bracket and fasteners with cross-laminated timber were tested, investigated and assessed for damage under seismic loading protocols for a hybrid application. An energy-based formulation according to Krätzig was applied to calculate the development of the damage index, and the resulting index was validated with visual observation. Six of the connections were modeled in OpenSees. For the modeling, a CUREE-10 parameter model was chosen to reproduce the test curves. The load-displacement results from both test and model were analyzed; the first method according to ASTM standards, where the envelope curve of the hysteretic results are considered and plotted in an equivalent energy elastic-plastic curve (EEEP). The second analyzing method used, was Krätzig’s damage accumulation model. Throughout all six combinations and both loading directions (parallel- and perpendicular-to-the-grain) a major difference was found in the analyzing methods. The EEEP curve roughly approximates the performance but with the damage accumulation method showed that analysis of the subsequent cycles is required to better reflect the empirical performance of the connections. To avoid the extensive destruction of a bracket type connection after completion of seismic loadings, a new approach was chosen. It was found that a tube connection can obtain comparably similar strength results as a conventional bracket connection. The computed mechanical properties of bracket-type and tube-type connections were compared and evaluated. The new tube connection showed great potential for future timber-steel hybrid structures and their connecting challenge. A total of 27 connection assemblies were tested under quasi-static monotonic and reversed cyclic loads. The tube connections showed two major differences when compared to traditional bracket connections: i) the completely linear elastic behaviour at the beginning, and ii) the continued load increase after yielding. Both phenomena are founded in the geometry of that connector effectively making the novel connector a very promising alternative.  iii Preface  A version of Chapter 2 has been published. Johannes Schneider, Karacabeyli E., Popovski M., Stiemer S. F., and Tesfamariam S. (2014) Damage assessment of connections used in cross laminated timber subjected to cyclic loads. Journal of Performance of Constructed Facilities, Vol.28, No.6: A4014008. I conducted all testing, analysis and was the lead author for the manuscript.  A version of Chapter 3 has been presented at the World Conference on Timber Engineering 2012 in Auckland, New Zealand and has been published in the conference proceedings. Johannes Schneider, Stiemer S.F., Tesfamariam S., Karacabeyli E., and Popovski M., (2012) Damage assessment of cross laminated timber connections subjected to earthquake loads. WCTE 2012 proceedings: Journal Monday: 398-406. I conducted all testing, analysis and was the lead author for the manuscript.  A version of Chapter 4 has been published. Johannes Schneider, Shen Y., Stiemer S.F., Tesfamariam S. (2015) Assessment and comparison of experimental and numerical model studies of cross-laminated timber connections. Construction and Building Materials, Vol. 77: 197-212. I conducted all the testing and wrote most of the manuscript. The sections on “Saws hysteretic model” and “Saws hysteretic model results” were originally drafted by Shen, Y.  A version of Chapter 5 has been submitted to ASCE Journal of Structural Engineering. Johannes Schneider, Tannert T., Stiemer S.F., Tesfamariam S. (2015) Assessment of experimental studies of a novel tube connector and conventional connections for Cross-Laminated Timber. I conducted all testing, processed the data, and I was the lead author for the manuscript.   ivTable of Contents  Abstract .................................................................................................................................... ii Preface ..................................................................................................................................... iii Table of Contents ................................................................................................................... iv List of Tables ........................................................................................................................ viii List of Figures .......................................................................................................................... x List of Symbols, Abbreviations or Other ......................................................................... xviii Acknowledgements ............................................................................................................... xx 1    Chapter: Introduction ...................................................................................................... 1 1.1 Research background and motivation ...................................................................... 1 1.2 Hybridization Types ................................................................................................. 2 1.3 Structural elements ................................................................................................... 6 1.3.1 Moment resisting frames ..................................................................................... 6 1.3.2 Braced frames ...................................................................................................... 6 1.3.3 Shear walls ........................................................................................................... 7 1.4 Wood-steel hybrid .................................................................................................... 8 1.5 Connections ............................................................................................................ 10 1.6 Damage to CLT connections.................................................................................. 14 1.7 Damage evaluation ................................................................................................. 15 1.8 Objectives and scope .............................................................................................. 17 1.9 Thesis structure ...................................................................................................... 18 2    Chapter: Damage assessment of connections used in cross laminated timber subject to cyclic loads ......................................................................................................................... 20 2.1 Framework for connection tests and quantification of damage ............................. 22 2.1.1 Connection tests ................................................................................................. 23 2.1.1.1 CLT samples................................................................................................ 23 2.1.1.2 Brackets and Fasteners ................................................................................ 24 2.1.1.3 Test Setup .................................................................................................... 28 2.1.1.4 Loading Protocol ......................................................................................... 29  v 2.2 Test results ............................................................................................................. 32 2.2.1 Hysteretic response ............................................................................................ 32 2.2.2 Failure Modes in CLT Connection Tests ........................................................... 37 2.3 Damage accumulation principle ............................................................................. 39 2.4 Comparison between calculated damage index and visual observation ................ 40 2.4.1 Analysis and comparison for tests parallel to the grain direction ...................... 41 2.4.2 Analysis and comparison for tests perpendicular to grain direction .................. 44 2.5 Preliminary Damage Prediction ............................................................................. 48 3    Chapter: Assessment and comparison of experimental and numerical model studies of Cross-laminated timber connections .............................................................................. 51 3.1 Experimental test and analytical models ................................................................ 56 3.1.1 Test outline ........................................................................................................ 56 3.1.2 SAWS hysteretic model ..................................................................................... 58 3.1.3 Analytical assessment Methods ......................................................................... 59 3.1.4 Ductility and equivalent energy elastic plastic curve ........................................ 59 3.1.5 Energy-based damage model ............................................................................. 60 3.2 Results and comparison of Test and Model ........................................................... 61 3.2.1 SAWS hysteretic model results ......................................................................... 61 3.2.2 Hysteretic Response .......................................................................................... 62 3.3 Performance assessment using the damage accumulation index ........................... 72 3.3.1 Results for connections parallel to the grain direction ...................................... 72 3.3.2 Results for connections perpendicular to the grain direction ............................ 75 3.4 Comparison of the model to the assessment method ............................................. 78 4    Chapter: Novel Steel Tube Connection for Hybrid Building Application ................ 80 4.1 Hybrid Structures ................................................................................................... 80 4.2 Ductility in Timber Structures ............................................................................... 81 4.3 Conventional Connections for CLT Walls ............................................................. 83 4.4 Steel tube connectors ............................................................................................. 85 4.4.1 Existing steel tube connectors ........................................................................... 85 4.4.2 Novel tube connector ......................................................................................... 86 4.4.3 Connector assembly ........................................................................................... 87 4.4.4 Materials ............................................................................................................ 89 4.4.5 Methods ............................................................................................................. 90  vi4.5 Results and discussion ........................................................................................... 91 4.5.1 Results of the static tests .................................................................................... 92 4.5.2 Results Cyclic Tests ........................................................................................... 96 4.6 Comparison to conventional connections .............................................................. 99 4.6.1 Failure modes .................................................................................................... 99 4.6.2 Static load-displacement curves ...................................................................... 100 4.6.3 Hysteretic curves ............................................................................................. 101 4.6.4 Equivalent energy elastic-plastic curves .......................................................... 103 4.6.5 Stiffness and Capacity ..................................................................................... 104 5    Chapter: Conclusion ..................................................................................................... 107 5.1 Main contributions ............................................................................................... 107 5.2 Recommendations for further research ................................................................ 111 Bibliography ........................................................................................................................ 113 Appendices ........................................................................................................................... 121 Test Results from all performed Connection Tests ........................................................... 121 Appendix A Bracket B with 10 Spiral Nails 4.2 mm x 89 mm ......................................... 122 A.1 Monotonic Test Results ................................................................................... 122 A.2 Cyclic Test Results .......................................................................................... 124 A.3 Pictures ............................................................................................................ 133 Appendix B Bracket A with 12 Ring Shank Nails 3.76 mm x 76 mm ............................. 135 B.1 Monotonic Test Results ................................................................................... 135 B.2 Cyclic Test Results .......................................................................................... 137 B.3 Pictures ............................................................................................................ 147 Appendix C Bracket A with 12 Ring Shank Nails 4.2 mm x 60 mm ............................... 150 C.1 Monotonic Test Results ................................................................................... 150 C.2 Cyclic Test ....................................................................................................... 152 C.3 Pictures ............................................................................................................ 162 Appendix D Bracket A with 9 Screws 5 mm x 90 mm ..................................................... 164 D.1 Monotonic Test Results ................................................................................... 164 D.2 Cyclic Test Results .......................................................................................... 166 D.3 Pictures ............................................................................................................ 176 Appendix E Bracket A with 18 Screws 4 mm x 70 mm ................................................... 178 E.1 Monotonic Test Results ................................................................................... 178  viiE.2 Cyclic Test Results .......................................................................................... 180 E.3 Pictures ............................................................................................................ 190 Appendix F Bracket A with 18 Spiral Nails 4.2 mm x 89 mm ......................................... 193 F.1 Monotonic Test Results ................................................................................... 193 F.2 Cyclic Test Results .......................................................................................... 195 F.3 Pictures ............................................................................................................ 208 Appendix G Tube with diameter 50.8 mm (2 inch) .......................................................... 213 G.1 Monotonic Test Results ................................................................................... 213 G.2 Cyclic Test Results .......................................................................................... 214 G.3 Pictures ............................................................................................................ 219 Appendix H Tube with diameter 76.2 mm (3 inch) .......................................................... 221 H.1 Monotonic Test Results ................................................................................... 221 H.2 Cyclic Test Results .......................................................................................... 222 H.3 Pictures ............................................................................................................ 228 Appendix I Tube with diameter 101.6 mm (4 inch) .......................................................... 231 I.1 Monotonic Test Results ....................................................................................... 231 I.2 Cyclic Test Results .............................................................................................. 232 I.3 Pictures ................................................................................................................. 237 Appendix J Damage assessment of cross-laminated timber connections and shear wall connections subjected to simulated earthquake loads .................................................................... 241 J.1 Introduction .......................................................................................................... 241 J.2 Connection test ..................................................................................................... 241 J.3 Shear wall test ...................................................................................................... 241 J.4 Loading protocol for shear wall tests ................................................................... 242 J.5 Experimental results and calibration of damage indices ...................................... 245 J.6 Failure modes of shear wall connections ............................................................. 245 J.7 Shear wall analysis ............................................................................................... 245 J.8 Damage prediction ............................................................................................... 251   viiiList of Tables  Table 2.1 Overview of tested samples and their features ................................................ 27 Table 2.2 Ductility ratio and elastic shear stiffness of the connections .......................... 33 Table 2.3 Summary of maximum forces and displacements ........................................... 36 Table 2.4 Minimum, Maximum and average values at 0.8Fmax ...................................... 42 Table 2.5 Minimum, maximum and average values at Fmax ........................................... 45 Table 2.6 Damage description for CLT connections (parallel-to-the-grain) ................... 48 Table 2.7 Damage description for CLT connections (perpendicular-to-the-grain) ......... 49 Table 2.8 Relationship between damage index and damage of connection test ............. 50 Table 3.1 Categories of damage principles ..................................................................... 53 Table 3.2 List of previous research on modelling wood connections ............................. 54 Table 3.3 Parameter estimation of SAWS model for monotonic connections tests ........ 61 Table 3.4 Parameter estimation of SAWS model for cyclic connections tests ............... 62 Table 3.5 Summary of ductility ratio, elastic shear stiffness, maximum forces, and  displacement at maximum force of the connections under cyclic loading  (Parallel-to-the-grain) for four test combinations ........................................... 66 Table 3.6 Summary of ductility ratio, elastic shear stiffness, maximum forces, and  displacement at maximum force of the connections under cyclic loading  (Parallel-to-the-grain) for two test combinations ............................................ 67 Table 3.7 Summary of ductility ratio, elastic shear stiffness, maximum forces, and  displacement at maximum force of the connections under cyclic loading  (Perpendicular-to-the-grain) for four test combinations ................................. 68 Table 3.8 Summary of ductility ratio, elastic shear stiffness, maximum forces, and  displacement at maximum force of the connections under cyclic loading  (Perpendicular-to-the-grain) for two test combinations .................................. 69 Table 3.9 Time of the loading protocol (parallel-to-the-grain) when damage index  reached D = 0.8 ............................................................................................... 73 Table 3.10 Amplitude of primary cycles for CUREE-protocol ........................................ 78 Table 3.11 Time of the loading protocol (perpendicular-to-the-grain) when damage  index reached D = 0.8 ..................................................................................... 78  ixTable 4.1 Classification of failure mode according to Smith et al. (2006) ..................... 82 Table 4.2 Test results of tube-type connections under static loading ............................. 93 Table 4.3 Test results of tube-type connections under cyclic loading ............................ 99 Table J.4.1 Shear wall tests used for analysis .................................................................. 244 Table J.8.1 Classification of Damage for CLT connections ............................................ 251   x List of Figures  Figure 1.1 Aerial view of Christchurch on February 22nd, 2011 right after the  earthquake (left);  aerial view of Christchurch in 2013 (right) ......................... 1 Figure 1.2 Structural hybridization systems (a) Tube in tube system (Green & Karsh,  2012); (b) Vertical mixed system (Craig, 2008); (c) Stiff core and light  wood frame system; (d) Steel frame with wood infill walls ............................. 5 Figure 1.3 Lateral force resisting systems: a) moment resisting frames,  b) braced frames, and  c) shear walls ................................................................ 8 Figure 1.4 Simpson StrongTie bracket 90×48×116, pull out failure (left);  BMF bracket 105 with rib, rupture failure (right) ........................................... 11 Figure 1.5 Connection with pins and insert steel plate ..................................................... 11 Figure 1.6 CLT panel connections .................................................................................... 12 Figure 1.7 European yield model with CLT modifications (Blass et al., 2007) ............... 14 Figure 2.1 Proposed timber-steel hybrid structure ........................................................... 21 Figure 2.2 Cross section of used CLT (board layers: 30 mm – 34 mm – 30 mm) ........... 24 Figure 2.3 Bracket A  (Simpson StrongTie) ..................................................................... 25 Figure 2.4 Bracket B (Simpson StrongTie) ...................................................................... 25 Figure 2.5 Applied fasteners with brackets ...................................................................... 26 Figure 2.6 Test Set-up for connection tests parallel to the grain ...................................... 28 Figure 2.7 Test Set-up for connection tests perpendicular to the grain with roller  in the back to prevent sample from bending ................................................... 29 Figure 2.8 Load-displacement curve, equivalent energy elastic-plastic (EEEP)  curve of a cyclic test and the reference deformation at 0.8Pmax ...................... 30 Figure 2.9 Cyclic schedule CUREE test protocol parallel-to-the-grain ........................... 31 Figure 2.10 Cyclic schedule CUREE test protocol perpendicular-to-the-grain ................. 31 Figure 2.11 Hysteretic response of connection test parallel-to-the-grain ........................... 34 Figure 2.12 Hysteretic response of connection test perpendicular-to-the-grain ................. 35 Figure 2.13 Different failure mode of connections (Figures show tests parallel  and perpendicular to the grain, therefore some brackets are oriented  horizontally) .................................................................................................... 38  xiFigure 2.14 Primary (PHC) and follower (FHC) half-cycles (Krätzig et al., 1989) ........... 40 Figure 2.15 Tearing through of connectors ........................................................................ 42 Figure 2.16 Cumulative damage index parallel-to-the-grain .............................................. 43 Figure 2.17 Combination parallel-to-the-grain (left), and combination  perpendicular-to-the-grain (right) ................................................................... 44 Figure 2.18 Cumulative damage index perpendicular-to-the-grain .................................... 46 Figure 2.19 Comparison between cumulative damage indices parallel-  and perpendicular-to-the-grain ........................................................................ 47 Figure 2.20 Relationship between visual damge and damage index .................................. 50 Figure 3.1 Proposed timber-steel hybrid structure (left), detail of connection (right) ..... 52 Figure 3.2 Flowchart to describe the procedure of damage accumulation assessment .... 56 Figure 3.3 Overview of tested connection combinations and test identification .............. 57 Figure 3.4 Hysteresis model for CLT connection ............................................................ 59 Figure 3.5 Equivalent energy elastic-plastic curve for test and SAWS  model parallel to the grain ............................................................................... 64 Figure 3.6 Equivalent energy elastic-plastic curve for test and SAWS model  perpendicular-to-the-grain ............................................................................... 65 Figure 3.7 Hysteretic response of connection tests and SAWS model  parallel-to-the-grain ......................................................................................... 70 Figure 3.8 Hysteretic response of connection tests and SAWS model  perpendicular-to-the-grain for two different bracket combinations ................ 71 Figure 3.9 Cumulative damage index parallel-to-the-grain .............................................. 74 Figure 3.10 Cumulative damage index perpendicular-to-the-grain .................................... 77 Figure 4.1 Timber-steel hybrid system (left) and observed damage of a  conventional hold-down connection under reversed cyclic load (right) ......... 81 Figure 4.2 Pull-out failure of nails in bracket-type connection (left);  deformed tube connection from Leijten (right) ............................................... 84 Figure 4.3 Considered novel tube connectors (left) and CLT panel (right) ..................... 87 Figure 4.4 Installation steps for tube connector ................................................................ 88 Figure 4.5 Test specimen with tube insert: side view (top)  and cross sectional top view (bottom) ............................................................. 89  xiiFigure 4.6 Test specimen in test frame ............................................................................. 91 Figure 4.7 Tested 2 inch tube connector beside CLT panel (left);  Tested tubes 2 inch, 3 inch, and 4 inch (right) ................................................ 94 Figure 4.8 Failure mode of a 76.2 mm tube connection under monotonic loading .......... 94 Figure 4.9 Schematic graph explaining the relation between measured load  and observed displacement .............................................................................. 95 Figure 4.10 Monotonic and cyclic load-deformation curve for 50.8 mm tube;  76.2 mm tube; 101.6 mm tube ......................................................................... 97 Figure 4.11 Load-displacement relation of static testing of tube-type and bracket  type connections ............................................................................................ 100 Figure 4.12 Hysteretic curves of tube connections and bracket connections ................... 102 Figure 4.13 Equivalent energy elastic-plastic curves under monotonic loading .............. 103 Figure 4.14 Equivalent energy elastic-plastic curves under cyclic loading ...................... 104 Figure 4.15 Overview of elastic shear stiffness Ke, yielding load Fy, maximum  load Fmax,  ultimate load Fu, and ductility ratio D under static loading ......... 105 Figure 4.16 Overview of elastic shear stiffness Ke, yielding load Fy, maximum  load Fmax,  ultimate load Fu, ductility ratio D, and equivalent viscous  damping ratio γRd under cyclic loading ......................................................... 106 Figure A.1.1 Monotonic test results for connection combination B-N-L-M ...................... 122 Figure A.1.2 Monotonic test results for connection combination B-N-P-M ...................... 123 Figure A.2.1 Cyclic test result for test combination CLT-B-N-L-C02 .............................. 124 Figure A.2.2 Cyclic test result for test combination CLT-B-N-L-C03 .............................. 125 Figure A.2.3 Cyclic test result for test combination CLT-B-N-L-C04 .............................. 126 Figure A.2.4 Cyclic test result for test combination CLT-B-N-L-C05 .............................. 127 Figure A.2.5 Cyclic test result for test combination CLT-B-N-P-C02 .............................. 128 Figure A.2.6 Cyclic test result for test combination CLT-B-N-P-C03 .............................. 129 Figure A.2.7 Cyclic test result for test combination CLT-B-N-P-C04 .............................. 130 Figure A.2.8 Cyclic test result for test combination CLT-B-N-P-C05 .............................. 131 Figure A.2.9 Cyclic test result for test combination CLT-B-N-P-C07 .............................. 132 Figure A.3.1 Bracket B under monotonic loading .............................................................. 133 Figure A.3.2 Pull-out failure of Bracket B under cyclic loading ....................................... 133  xiiiFigure A.3.3 Bracket B in test set-up for testing perpendicular to the grain ...................... 134 Figure A.3.4 Fracture failure of Bracket B under cyclic loading:  tension loading (left), compression loading (right) ....................................... 134 Figure B.1.1 Monotonic test results for connection combination A-R-L-M ...................... 135 Figure B.1.2 Monotonic test results for connection combination A-R-P-M ...................... 136 Figure B.2.1 Cyclic test result for test combination CLT-A-R-L-C04 .............................. 137 Figure B.2.2 Cyclic test result for test combination CLT-A-R-L-C05 .............................. 138 Figure B.2.3 Cyclic test result for test combination CLT-A-R-L-C06 .............................. 139 Figure B.2.4 Cyclic test result for test combination CLT-A-R-L-C07 .............................. 140 Figure B.2.5 Cyclic test result for test combination CLT-A-R-L-C08 .............................. 141 Figure B.2.6 Cyclic test result for test combination CLT-A-R-P-C01 .............................. 142 Figure B.2.7 Cyclic test result for test combination CLT-A-R-P-C02 .............................. 143 Figure B.2.8 Cyclic test result for test combination CLT-A-R-P-C06 .............................. 144 Figure B.2.9 Cyclic test result for test combination CLT-A-R-P-C07 .............................. 145 Figure B.2.10 Cyclic test result for test combination CLT-A-R-P-C08 .............................. 146 Figure B.3.1 Bracket A with 12 Ring Shank Nails under monotonic loading ................... 147 Figure B.3.2 Bracket A under cyclic loading: one nail failed in shear,  one nail failed in pull-out .............................................................................. 147 Figure B.3.3 Four stages of a cyclic loading with Bracket A ............................................ 148 Figure B.3.4 Bracket A under cyclic loading perpendicular to the grain ........................... 149 Figure C.1.1 Monotonic test results for connection combination A-r-L-M ....................... 150 Figure C.1.2 Monotonic test results for connection combination A-r-P-M ....................... 151 Figure C.2.1 Cyclic test result for test combination CLT-A-r-L-C01................................ 152 Figure C.2.2 Cyclic test result for test combination CLT-A-r-L-C02................................ 153 Figure C.2.3 Cyclic test result for test combination CLT-A-r-L-C03................................ 154 Figure C.2.4 Cyclic test result for test combination CLT-A-r-L-C04................................ 155 Figure C.2.5 Cyclic test result for test combination CLT-A-r-L-C05................................ 156 Figure C.2.6 Cyclic test result for test combination CLT-A-r-P-C03 ................................ 157 Figure C.2.7 Cyclic test result for test combination CLT-A-r-P-C04 ................................ 158 Figure C.2.8 Cyclic test result for test combination CLT-A-r-P-C05 ................................ 159 Figure C.2.9 Cyclic test result for test combination CLT-A-r-P-C06 ................................ 160  xiv Figure C.2.10 Cyclic test result for test combination CLT-A-r-P-C07 ................................ 161 Figure C.3.1 Front view of Bracket A with short Rink Shank Nails ................................. 162 Figure C.3.2 Pull-out failure of Bracket A with small Ring Shank Nails  under cyclic loading ...................................................................................... 162 Figure C.3.3 Front view of test set-up perpendicular to the grain ...................................... 163 Figure C.3.4 Destructed CLT panel after test perpendicular to the grain .......................... 163 Figure D.1.1 Monotonic test results for connection combination A-S-L-M ...................... 164 Figure D.1.2 Monotonic test results for connection combination A-S-P-M ...................... 165 Figure D.2.1 Cyclic test result for test combination CLT-A-S-L-C07 ............................... 166 Figure D.2.2 Cyclic test result for test combination CLT-A-S-L-C08 ............................... 167 Figure D.2.3 Cyclic test result for test combination CLT-A-S-L-C09 ............................... 168 Figure D.2.4 Cyclic test result for test combination CLT-A-S-L-C10 ............................... 169 Figure D.2.5 Cyclic test result for test combination CLT-A-S-L-C ................................... 170 Figure D.2.6 Cyclic test result for test combination CLT-A-S-P-C05 ............................... 171 Figure D.2.7 Cyclic test result for test combination CLT-A-S-P-C06 ............................... 172 Figure D.2.8 Cyclic test result for test combination CLT-A-S-P-C08 ............................... 173 Figure D.2.9 Cyclic test result for test combination CLT-A-S-P-C09 ............................... 174 Figure D.2.10 Cyclic test result for test combination CLT-A-S-P-C10 ............................... 175 Figure D.3.1 Bracket A with 9 Screws 5 x 90 mm under monotonic loading ................... 176 Figure D.3.2 Deformed connector Bracket A under cyclic loading ................................... 176 Figure D.3.3 Cracked CLT panel under cyclic loading perpendicular to the grain ........... 177 Figure D.3.4 Shear failure of all nine screws under cyclic loading .................................... 177 Figure E.1.1 Monotonic test results for connection combination A-s-L-M ....................... 178 Figure E.1.2 Monotonic test results for connection combination A-S-L-M ...................... 179 Figure E.2.1 Cyclic test result for test combination CLT-A-s-L-C04 ............................... 180 Figure E.2.2 Cyclic test result for test combination CLT-A-s-L-C05 ............................... 181 Figure E.2.3 Cyclic test result for test combination CLT-A-s-L-C06 ............................... 182 Figure E.2.4 Cyclic test result for test combination CLT-A-s-L-C07 ............................... 183 Figure E.2.5 Cyclic test result for test combination CLT-A-s-L-C08 ............................... 184 Figure E.2.6 Cyclic test result for test combination CLT-A-s-P-C03................................ 185 Figure E.2.7 Cyclic test result for test combination CLT-A-s-P-C04................................ 186  xv Figure E.2.8 Cyclic test result for test combination CLT-A-s-P-C07................................ 187 Figure E.2.9 Cyclic test result for test combination CLT-A-s-P-C08................................ 188 Figure E.2.10 Cyclic test result for test combination CLT-A-s-P-C09................................ 189 Figure E.3.1 Bracket A with 18 screws 4 x 70 mm monotonic loaded (front view) ......... 190 Figure E.3.2 Bracket A with 18 screws 4 x 70 mm monotonic loaded (side view) ........... 190 Figure E.3.3 Series of four stages of Bracket A (cyclic loading) parallel-to-the-grain ..... 191 Figure E.3.4 Front view of Bracket A with screws 4 x 70 mm under cyclic  loading perpenicular to the grain ................................................................... 192 Figure E.3.5 Destructed CLT panel under cyclic loading perpendicular to the grain ....... 192 Figure F.1.1 Monotonic test results for connection combination A-N-L-M ..................... 193 Figure F.1.2 Monotonic test results for connection combination A-N-P-M ...................... 194 Figure F.2.1 Cyclic test result for test combination CLT-A-N-L-C01 .............................. 195 Figure F.2.2 Cyclic test result for test combination CLT-A-N-L-C02 .............................. 196 Figure F.2.3 Cyclic test result for test combination CLT-A-N-L-C03 .............................. 197 Figure F.2.4 Cyclic test result for test combination CLT-A-N-L-C04 .............................. 198 Figure F.2.5 Cyclic test result for test combination CLT-A-N-L-C05 .............................. 199 Figure F.2.6 Cyclic test result for test combination CLT-A-N-L-C06 .............................. 200 Figure F.2.7 Cyclic test result for test combination CLT-A-N-L-C07 .............................. 201 Figure F.2.8 Cyclic test result for test combination CLT-A-N-P-C02 .............................. 202 Figure F.2.9 Cyclic test result for test combination CLT-A-N-P-C03 .............................. 203 Figure F.2.10 Cyclic test result for test combination CLT-A-N-P-C04 .............................. 204 Figure F.2.11 Cyclic test result for test combination CLT-A-N-P-C05 .............................. 205 Figure F.2.12 Cyclic test result for test combination CLT-A-N-P-C06 .............................. 206 Figure F.2.13 Cyclic test result for test combination CLT-A-N-P-C07 .............................. 207 Figure F.3.1 Front view of Bracket A with 18 Spiral Nails under  monotonic loading, parallel to the grain ........................................................ 208 Figure F.3.2 Pull-out failure of Spiral Nails in combination with Bracket A .................... 208 Figure F.3.3 Failed connection with Bracket A; Pull-out failure (monotonic loading) ..... 209 Figure F.3.4 CLT panel after cyclic loading perpendicular to the grain;  combined pull-out failure with shear failure of fasteners ............................. 209 Figure F.3.5 Stages 1 to 4 of Bracket A with Spiral Nails under cyclic loading ............... 210  xvi Figure F.3.6 Stages 5 to 8 of Bracket A with Spiral Nails under cyclic loading ............... 211 Figure F.3.7 Bracket A with 18 Spiral Nails under monotonic loading  perpendicular to the grain .............................................................................. 212 Figure G.1.1 Monotonic test results for connection combination T2-1-M ........................ 213 Figure G.2.1 Cyclic test result for test combination CLT-T2-1-C01 ................................. 214 Figure G.2.2 Cyclic test result for test combination CLT-T2-1-C02 ................................. 215 Figure G.2.3 Cyclic test result for test combination CLT-T2-1-C03 ................................. 216 Figure G.2.4 Cyclic test result for test combination CLT-T2-1-C04 ................................. 217 Figure G.2.5 Cyclic test result for test combination CLT-T2-1-C05 ................................. 218 Figure G.3.1 Test set-up of T2 ........................................................................................... 219 Figure G.3.2 Loading stages of T2; no visible destruction of CLT panel .......................... 219 Figure G.3.3 Destroyed Tube T2; CLT panel without any failure ..................................... 220 Figure H.1.1 Monotonic test results for connection combination T3-1-M ........................ 221 Figure H.2.1 Cyclic test result for test combination CLT-T3-1-C01 ................................. 222 Figure H.2.2 Cyclic test result for test combination CLT-T3-1-C02 ................................. 223 Figure H.2.3 Cyclic test result for test combination CLT-T3-1-C03 ................................. 224 Figure H.2.4 Cyclic test result for test combination CLT-T3-1-C04 ................................. 225 Figure H.2.5 Cyclic test result for test combination CLT-T3-1-C05 ................................. 226 Figure H.2.6 Cyclic test result for test combination CLT-T3-1-C06 ................................. 227 Figure H.3.1 Test set-up for T3 under cyclic loading ........................................................ 228 Figure H.3.2 Deformed tube T3 under monotonic loading ................................................ 228 Figure H.3.3 Close up of failed tube T3 ............................................................................. 229 Figure H.3.4 Deformed tube T3 under cyclic loading ........................................................ 230 Figure I.1.1 Monotonic test results for connection combination T4-1-M ........................ 231 Figure I.2.1 Cyclic test result for test combination CLT-T4-1-C05 ................................. 232 Figure I.2.2 Cyclic test result for test combination CLT-T4-1-C02 ................................. 233 Figure I.2.3 Cyclic test result for test combination CLT-T4-1-C03 ................................. 234 Figure I.2.4 Cyclic test result for test combination CLT-T4-1-C04 ................................. 235 Figure I.2.5 Cyclic test result for test combination CLT-T4-1-C05 ................................. 236 Figure I.3.1 Test set-up tube T4 ........................................................................................ 237 Figure I.3.2 Deformed tube T4 after cyclic loading ......................................................... 237  xvii Figure I.3.3 Close-up of tested tube T4 ............................................................................ 238 Figure I.3.4  MTS test machine set-up for tube connections with base steel beam ........... 239 Figure I.3.5 Tubes T2, T3, and T4 after loaded with cyclic loading protocol .................. 240 Figure J.4.1 Isometric drawing of the shear wall test set-up ............................................. 243 Figure J.7.1 Failure modes observed in CLT shear wall tests ........................................... 246 Figure J.7.2 Wall results for group 1 (2.3m x 2.3) ............................................................ 247 Figure J.7.3 Wall results for group 3 (2.3m x2.3m) .......................................................... 248 Figure J.7.4 Wall results for group 5 (2.3m x 3.45m) ....................................................... 249 Figure J.7.5 Wall results for group 6 (3.45m x 2.3m) ....................................................... 250   xviiiList of Symbols, Abbreviations or Other  A = area under the curve from zero to ultimate displacement ∆u a = Slope of regression; b,c = Locally defined constants; CLT = cross-laminated timber COV = coefficient of variation D-  = damage in negative cycle d = bolt diameter D = Damage index; D+  = damage in positive cycle D+D-  = interaction of D+ and D- De  = energy-related damage index E = Energy; e = y intercept; E+f  = energy in a monotonic test to failure E+i  = energy in a FHC E+p,i  = energy in a PHC EEEP = Energy equivalent elastic-plastic curve Ep,i = available potential engergy f = Failure; F = Force; F0 = intercept strength fh,1,1 = characteristic strength of wood layer 1 fh,1,k  = characteristic strength of wood member FHC = Follower half-cycle;  Fi  = force in i-th cycle FI = intercept strength for the pinching part (FI > 0) Fy  = force at yielding Ke  = elastic shear stiffness KP = degrading stiffness m = Maximum; My,k = characteristic yield moment  xix NCR = Normalized cumulative; p = panel perimeter nail spacing;  p = Primary half-cycle; PHC = Primary half-cycle Rotation; Ppeak = maximum measured load Pu = Ultimate load at 0.8 Ppeak Pyield = Load at yielding point R1 = stiffness ratio of the asymptotic line to the backbone curve of (0 < R1 < 1.0) R2 = stiffness ratio on the descending segment of the envelope curve (R2 < 0) R3 = stiffness ratio of the unloading segment off the envelope curve (R3<1) R4 = stiffness ratio of the pinching part of the hysteretic curve (R4 > 0) S0  = initial stiffness of the hysteretic curve (S0 > 0) t1 = entire thickness of wood member u = ultimate; xun = last unloading displacement off the envelop curve β1,1,2 = strength ratio between layer 1(fh,1,k) and layer 2(fh,2,k), δ = Deformation; ∆ = Displacement; ∆e = corresponding displacement at 0.4 Ppeak ∆i  = displacement in i-th cycle ∆ref = reference displacement to calculate loading protocol ∆peak = displacement at maximum load ∆u = displacement at ultimate load ∆yield = displacement at yield load θ = Rotation; µ = Ductility ratio; ϕ = Curvature; ψ = Calibration parameter ψ = thickness ration between layer 1(t1) and layer 2(t2) µ = mean value σ = standard derivation νu  = ultimate slip  xx Acknowledgements  This thesis would not have been possible without the financial support of several organizations. This research was supported through funding to the NSERC Strategic Network on Innovative Wood Products and Building Systems - NEWBuildS and the Steel Structures Education Foundation (SSEF). First, I would like to express my enduring gratitude to my Ph.D. supervisors for giving me the opportunity to be part of their research. I greatly appreciate, that I was always encouraged and supported to participate in national and international conferences, presenting the outcomes of the research work. My main supervisor Dr. Solomon Tesfamariam guided me through my Ph.D. research with great dedication. His advice and ideas have contributed greatly to the outcomes of this research project. His encouragement, advice and patience were highly appreciated. I would like to express further sincere gratitude to my co-supervisor Dr. Siegfried F. Stiemer. Dr. Stiemer provided me with very valuable advice and suggestions for my experimental research work and shared with me his ideas and his rich experiences.I would like to thank my committee member Dr. Thomas Tannert for his support and advice with his great expertise in timber engineering through my time at UBC. Thanks are also expressed to the staff of the machine shop of University of British Columbia for their help for preparations of the experimental tests. I would like to express profound acknowledgement to FPInnovations Vancouver and in particular to Dr. Marjan Popovski and Erol Karacabeli for their limitless support in providing me space and the opportunity to perform all my experimental work in their labs, their kindness and professional guidance. My very deepest gratitude to my nearest: my parents and my brother, who supported me all the time during my studies, and to my dear friend Delaney P. Boyd who was always there for me when I needed them.    11    Chapter: Introduction  1.1 Research background and motivation  Earthquakes are a natural disaster which cannot be avoided. Earthquakes occur all over the world with high prone zones along the fault lines. Structural engineering focuses on superior building performance in a seismic event. The primary objective of seismic design is to provide an acceptable level of safety for building occupants and the general public as the building responds to strong ground motions: i.e. collapse prevention and minimize loss of life (NRCC, 2011). Seismic events in the past years have shown a tremendous impact of earthquakes on the post-disaster management. On February 22nd, 2011 Christchurch in New Zealand was hit by a 6.3-magnitude earthquake. Most buildings did not collapse in the earthquake (Figure 1.1 left). The seismic performance met the structural requirements, as described above. However, the aerial view taken two years later depicts the real impact of that earthquake. A big number of buildings had to be taken down due to excessive and irreparable damage to the structures, mostly concrete and masonry buildings (Figure 1.1 right).   Figure 1.1    Aerial view of Christchurch on February 22nd, 2011 right after the earthquake (left);  aerial view of Christchurch in 2013 (right)   2The question arises, how such post-earthquake scenarios can be improved or avoided. A lot of research has been done on reinforcement of concrete structures after seismic impacts (Tsonos, 2008). Another alternative could be using more wood, in particular mass-timber or hybrids involving mass-timber as structural systems for mid-rise buildings. The structural use of mass-timber (e.g. CLT) is gaining popularity in North America. One reason is the height limitation imposed on light-frame timber structures to six storeys in some jurisdictions and four storeys in others (e.g. (NRCC, 2011; BCBC, 2010).  To further extend the performance of wood structures, steel can be incorporated in the design (Stiemer et al., 2012; Dickof et al., 2012). These wood-steel hybrid structures can be economical, safe and aesthetically pleasing (Leckie, 2008), however, to best realize the opportunities of hybrid construction, new techniques in design and construction are needed (Taranath, 1998). A major challenge in wood-steel hybrid construction lies in connecting the wood and steel elements (Schneider et al., 2012). This research will address two primary objectives: 1. Analyze and assess conventional connecting methods for steel-framed structures that incorporate shear walls of engineered wood, and 2. Provide an alternative connecting solution for timber-steel hybrid shear walls.  1.2 Hybridization Types  Hybridization of timber and steel can be found from material hybridization to hybridization of entire structures (system hybridization). Timber-steel hybrid structural systems are not yet covered by any structural design standard that guide the general implementation. For timber light-framed structures, which are often used for residential buildings with four or less levels, the National Association of Home Builders Research Center (NAHB, 2003) has developed a builder’s guide to hybrid wood and steel connection details. Timber-steel hybrids are not very common, but based on existing structures the two materials can be implemented at different levels, such as: •  Material, where both materials are combined in a structural element,  3•  System, involving material assemblies to create a structural system, and •  Structure, where the entire building is constructed of both materials. For this research, the structure hybridization is considered and will be discussed in the following paragraph. The next level of hybridization is the Structure Hybridization. There are several structural systems which can be applied to timber-steel hybrid structures. Structural hybrid systems can be classified into the following categories (Taranath, 1998). a) Tube systems, Figure 1.2a (e.g. Green et al. 2012) b) Vertically mixed systems, Figure 1.2b (e.g. Van de Lindt et al., 2012) c) Stiff core and wood light-frame systems, Figure 1.2c (e.g. Taranath, 1998) d) Steel frame with wood infill wall configuration, Figure 1.2d (e.g. Stiemer et al., 2012; Moore, 2000) Tube systems (Figure 1.2a) consist of an exterior enclosure wall system and an interior core. The tube type walls are made of engineered wood products such as CLT or laminated veneer lumber. A steel structure connects the wall panels between each other as well as creates a connection between the inner and outer tube. Green et al. analyzed in their case study of this system for tall buildings in regards to structural, fire, and energy efficiency. The FEM analysis showed great seismic performance up to fifteen levels. Beyond fifteen levels requires additional lateral load resisting systems (Green et al., 2012). To provide open space and avoiding big timber dimensions in the main level a vertical mixed system is often applied where the main level is built in steel or concrete (Figure 1.2b). All other levels above are added in timber light-framed structures. Within extensive studies a 7-storey building was tested on a shake table in Japan (Van de Lindt et al., 2012). Besides regular timber frame walls a number of Midply walls were installed. In the entire building, no structural wood failure occurred. Damage was observed only in nonstructural drywall elements. Another structural combination is the stiff core and attached light frames (Figure 1.2c). The core can either be built up of solid shear walls or braced frames (Taranath, 1998). The envelope of the building is created by light timber curtain walls. A steel frame structure with timber infill shear walls and timber diaphragms is the last one in the classification of structural hybrid systems (Figure 1.2d) (Stiemer et al., 2012). The  4steel frame structure carries the gravity load whereas the infill shear walls and diaphragms are responsible for the lateral loads. A good example for that type of structure is the 12-storey Scotia Place in Auckland/ New Zealand which was completed in 2000 (Moore, 2000). The main objective of the design was to find a very cost-effective structural system that would meet all code and building functional requirements. Aside from the structural system, fire performance, floor vibration, and acoustics were the main factors governing the design. The primary structure for the lateral load consists of concentrically braced steel frames. The glue laminated wood diaphragms transfer the load to the steel beams and from there to the concentrically braced steel frames with tubular steel columns. The Structure level is the highest level of hybridization. The structural parts are constructed by using timber and steel members and the internal forces on the structure are shared between the elements according to their best performance.  A feasibility study on a 14-storey building in New Zealand showed significant advantages of building structures consisting of steel moment resisting frame and composite timber long span floor joist & plywood flooring (Banks, 2006). This 14-storey building utilizes a perimeter moment resisting frame of structural steel with a structural steel internal gravity post & beam system. The flooring system which supports gravity loads and acts as a diaphragm to transfer lateral loads to the steel frames was constructed from timber HYBEAM joists and plywood flooring. HYBEAM is an all wood I-beam.  The concept of using a timber-steel hybrid structure required a number of unique design considerations. As a result of the significant reduced overall building weight, the lateral design was no longer governed by seismic forces (Moore, 2000). The design wind load was governed the structural design. There are also construction benefits with respect to the possibility of prefabricating lightweight sections of floor panels which could be lifted into place with less crane capacity than would be required for heavier materials.   5a)  b)   c)  d)  Figure 1.2    Structural hybridization systems (a) Tube in tube system (Green & Karsh, 2012); (b) Vertical mixed system (Craig, 2008); (c) Stiff core and light wood frame system; (d) Steel frame with wood infill walls   6 1.3 Structural elements  In buildings, individual structural elements are dedicated for specific functions. Gravity loads such as dead load, live load, or snow are taken by beams and columns. Lateral loads generated by wind or seismic events need a different system to guarantee structural integrity. There are three different approaches to provide enough stiffness in lateral direction: a) Moment resisting frame b) Braced frame c) Shear walls In the following subsection the three mentioned principles will be discussed.  1.3.1 Moment resisting frames  A moment resisting frame (MRF) (Figure 1.3a) is a common lateral load resisting system for steel structures (Schreibmair et al., 2010; Fragiacomo et al., 2010). It offers a great flexibility of space which is often required for retail space or open offices. The high strength of steel allows creating a very strong connection in a relatively small space at the corners of a moment frame. Both welded and bolted connections between columns and beams can be manufactured off-site or on-site in a precise manner. Steel with its high strength provides ideal properties resulting in the desirable smaller member and beam cross sections.  1.3.2 Braced frames  Using braced frames (Figure 1.3b) is an inexpensive way to provide lateral stability to a building. Concentrically braced frames are often used where large open spaces are required (Popovski et al., 2003). The bracings can be chosen as pure tension members or tension and compression members. The general failure modes of steel bracings occur as tensile yielding or buckling of the brace under compression (Yoo et al., 2008). Wooden bracings show the same failure  7modes, where the tensile yielding is achieved within the connectors between the timber bracing element and the connecting steel plates at the ends (Popovski et al., 2003). Wood bracings are better in compression than tension. The compression connection in wood is easier to manufacture and the maximum load capacity is higher than a tension connection (Li et al., 2009).  1.3.3 Shear walls  Shear walls (Figure 1.3c) are substantial elements for buildings. Shear walls provide lateral stability as well as collect forces generated by wind and seismic events. Steel structures achieve their stiffness either through braced frames, stiff cores or MRF. For a hybrid system, wood shear walls should be utilized to guarantee lateral stability while steel columns and beams look after gravity loads and transfer them to the base.  Wood-frame construction is the most common structural type in North America for single-family houses and low-rise multi-family dwellings, constituting around 90 % of all residential housing. Over the years, a number of research programs have been accomplished to study lateral stability under seismic loads. Van de Lindt (2006) did an extensive research on damage and damage accumulation of regular wood-frame shear walls where the stud spacing is 16 inches and the plywood sheathing is attached to one side. After individual cyclic test programs, different failure modes were identified and evaluated according to a combined damage accumulation factor. It was found that the connection between studs and sheathing provides the most ductility, which makes a timber-framed shear wall an excellent shear wall type although the capacity is limited (Van de Lindt et al., 2006). A modified wood-frame shear wall was developed and patented by Forintek and UBC (Varoglu et al., 2007). Varoglu (2007) placed the sheathing in the middle with studs on either side of the plywood panel. All nails were driven in alternating from both sides. The product was called Midply and test results demonstrated that Midply walls have superior survival characteristics under earthquake loading, and have a dynamic load-carrying capacity of more than two and a half times that of comparable standard shear walls. The nails between sheathing and studs are exposed to double shear, which makes these light-frame walls much stiffer compared to conventional walls (Varoglu et al., 2007). Dickof et al. (2012) compared different wood infill  8shear walls in their research. It was found that even a stiffer Midply wall may not add sufficient stiffness for lateral stability to a steel primary structure.   Figure 1.3    Lateral force resisting systems: a) moment resisting frames, b) braced frames, and  c) shear walls  1.4 Wood-steel hybrid  Wood-steel hybrid construction has several benefits: • a high degree of prefabrication;  • optimization of construction height;  • improvement of earthquake resistance through weight reduction; and  In contrast to wood-concrete composites, wood-steel hybrid structures have the advantage of being light and fast to erect. Hybridization of different materials is an emerging concept, which started with steel and concrete for bridge decking (Taranath, 1998). Hybrid structures utilize the best properties of each material and combine them to form a strong composition. In general, the different material property compatibilities cause a challenge in  9the hybrid design (temperature, humidity and anisotropy of wood are major effects to be considered), and in particular the connection design. Moore (2000) reported the application of a hybrid 12-storey building where a concentrically braced frame was used for lateral load resistance with a glulam timber floor slab. Projects applying wood-steel hybrids are: Karikobozu Bridge in the Miyazaki region, Japan (Imura et al., 2004), Vihantasalmi Bridge in Finland (Thelandersson et al., 2004), Bochu-Bashi Bridge in Akita region, Japan (2001), Brentwood Sky Train station Vancouver, Canada (2002), and Airport of Prince George, Canada (2004). Most of these applications were for infrastructure and commercial buildings. For the next generation of wood-steel hybrid structures it would be desirable to apply this system more in residential multi-storey buildings because the advantages mentioned earlier are obvious. Seismic design consideration is one of the major concerns in areas with moderate to high seismic hazard. The 2010 BC Building Code (BCBC, 2010), for example, has increased the limit for mid-rise wood-frame residential buildings up to six levels. To overcome this height limitation while using significant amounts of wood products, a timber-steel hybrid system can be a solution to that problem. To make a timber-steel hybrid structure attractive to engineers, it is important to develop and provide a robust hybrid design guideline. Nishiyama et al. (2004) have outlined an earthquake-resistant design guideline for hybrid structures (steel and concrete) using the working stress design method. The seismic design consideration utilizes capacity design procedure, where the beam is expected to dissipate the demand. This entails utilization of weak beam–strong column failure mechanism in designed buildings. Pang and Rosowsky (2007) have developed a displacement procedure for performance-based seismic design of mid-rise wood-framed structures. In regions of high seismicity, however, the progress of hybrid structures has been lagging, due to difficulty with joints detailing (Taranath, 1998). With the development of new engineered wood products hybridization in combination with wood became more popular in recent years. Various studies are reported on connection detailing for wood-steel structures. Studies showed that it is possible to build semi-rigid frames with glued-in rods without obvious bracing and to have better flexibility in architectural design of the building (Vašek et al., 2006); Laminated veneer lumber (LVL) subassemblies confirmed the enhanced performance of the jointed ductile or hybrid connections. (Palermo et al., 2006); Murty  10(2008) stated and proved in a test series that small diameter steel tube fasteners are an effective means of achieving strong, yet ductile structural wood connections; Connections with thin dowels can reach high load-carrying efficiency. For joints with multiple dowel type fasteners, ductile behaviour is typically only possible if the failure occurs after significant plastic deformation of the steel dowels (Mischler et al., 2000). Keller and Gürtler (2006) have performed experimental work on steel-wood hybrid beams. Limited amount of research results are available on the performance of hybrid steel-wood system response and performance.  According to the present BC building code (BCBC, 2006) multi-storey wood frame buildings may not exceed six storeys. Unlike strictly wood structures, innovative timber-steel hybrid building systems enable flexibility in design (e.g., greater than six storeys) beyond the limitations of wood alone. The steel part of the structure provides ductility and strength for beams and columns, while timber elements address lateral stiffness and enable efficiency in installation and load transfer for flooring. Timber and steel are quite different materials with different properties. The question raises how to connect them efficiently and meet all requirements in terms of installation, flexibility, and stiffness.  1.5 Connections  For timber-steel hybrid structures, the connections play a significant role. The connections between the two materials must meet several functional requirements. Such elements should be stiff enough to resist lateral and vertical forces while they remain sufficiently flexible to dissipate energy in order to perform well under earthquake loading (Figure 1.4). Aspects of accessibility and ease of manufacturing and installation, construction friendliness as well as fire performance need to be considered for a good connection (Leijten, 1998).  11  Figure 1.4    Simpson StrongTie bracket 90×48×116, pull out failure (left); BMF bracket 105 with rib, rupture failure (right)   Figure 1.5    Connection with pins and insert steel plate  The commercial market provides only few solutions for timber-steel hybrid connections, although steel connectors for pure wood-to-wood connections became very popular, because they are easy to install and very strong. It was common to use bar-shaped elements as connectors such as pins with insert steel plates (Figure 1.5). With introduction of mass wood products such as laminated strand lumber (LSL), laminated veneer lumber (LVL)  12and CLT connections required even more attention, because of the larger size of the panel-shaped elements (Figure 1.6).   Figure 1.6    CLT panel connections  The wood panels were very stiff and can not absorb much energy within the panel (Ceccotti et al., 2006). Thus, the energy dissipation must be achieved by the connections. The resulting connection design was proprietary. It provided good ductility and energy dissipating performance. However, the damage to the wood panel was often significant (Ceccotti et al., 2006). Similar systems with CLT wall panels and proprietary connections had been extensively tested in Slovenia (Dujic et al., 2004; Fragiacomo et al., 2011) and Canada (Schneider, 2009) exhibiting similar sacrificial failure modes. Beside wood crushing and shear failure of the fasteners, the most common and desirable failure mode was a pull-out failure of the fasteners. Pull-out failure provides ductility in the connection and can be controlled the best. Connections tested by Schneider showed a ductility ratio between 3 and 6. Depending on the fastener type, the maximum capacity varied between 26 kN and 60 kN.  The ratio was determined as displacement at failure over displacement at yielding. A different method of connection involves inserted steel plates. Various methods of inserting and fastening those inserts were studied. An interesting solution was the use of low yield strength steel tubes (Murty et al., 2011) replacing solid drift pins in order to provide more ductility within the connection. It was found that connection forces were distributed  13evenly within the group of steel tubes. The hollow steel tube provides the desired ductility and allowed larger displacements while dissipating energy (Figure 4.8). The tested samples showed an average ductility ratio in the order of 6. Steel tubes in size 6.35 mm (¼"), 9.52 mm (⅜"), and 12.7 mm (½") diameter were used in set-ups of 4 fasteners per connection and the peak loads were achieved respectively to 24.0 kN, 45.1 kN, and 64.4 kN. Steel tube groups were also used in wood-to-wood connections when the surfaces of the connecting members were reinforced with densified veneer wood (DVW). DVW has a density of 1300 kg/m3 and in combination with the wood member, it improves the bearing resistance of the outer wood fibers. The steel tube was inserted in an oversized hole (diameter enlarged by 1-2 mm) and expanded to a perfect fit. This causes pre-stressing of the surrounding timber and increases the stiffness of the joint (Leijten, 1998). Such a method provides enough stiffness that moment resisting timber frames can be designed. The maximum capacity for 18mm steel tube was measured to be 35.8 kN. The post yield part of the load-displacement curve was completely developed, which provided the sufficient ductility.  Leijten et al. (2011) also studied the lateral stiffness of CLT infill wall panels, in particular how insert steel plates with slotted holes affect the behavior of the load path. It was suggested to oversize the holes perpendicular to the beam/column in order to restrict the transfer of loads to the direction parallel to the edge of the infill wall panel.  When CLT became more popular, conventional fasteners and drift pins and their bearing resistance behavior in the individual layers of lumber were studied. Blass and Uibel (2007) derived a formula for connection design in CLT considering the fastener diameter, wood density and loading angle relative to the grain direction of the surface lamina. Empirical behavior models were developed based on test results. Johansen’s theory (Johansen, 1949) contributed to establish the embedment strength of the dowel type fasteners for lateral loads (Blass et al., 2007). Johansen’s theory (also called European Yield Model) has different failure modes based on thickness of the involved members. Figure 1.7 illustrates exemplarily the modification done by Blass for one failure mode. In a newer study, reinforced timber joints with self-drilling screws, densified wood, and punched metal plate additions were used to avoid wood failure thus increasing ductility. It was found that self-drilling screws are a very efficient and cost-effective method (Blass et al., 2012).  14     Where fh,1,k = characteristic strength of wood member, t1 = entire thickness of wood member, d = bolt diameter, My,k = characteristic yield moment, fh,1,1 = characteristic strength of wood layer 1, β1,1,2 = strength ratio between layer 1(fh,1,k) and layer 2(fh,2,k), ψ = thickness ration between layer 1(t1) and layer 2(t2) Figure 1.7    European yield model with CLT modifications (Blass et al., 2007)  1.6 Damage to CLT connections  CLT is a relatively young product on the market, and therefore, has only quite recently been subject to tests of its parameters and characteristics. Institutions in Europe and North America have conducted CLT tests under various conditions (e.g., Tree and Timber Institute in San Michele all’Adige, Italy (Ceccotti et al., 2006); University of Ljublijana, Slovenia (Dujic et al., 2004); University of Karlsruhe, Germany (Blass et al., 2007); and FPInnovations, Vancouver, Canada (Popovski et al., 2010 and Schneider, 2009). The primary focus for all tests to date has been on the connections under monotonic and cyclic loading, the latter to simulate seismic events. Primarily commercial anchor systems have been investigated, although some custom-made anchors were used in order to obtain a higher strength (Latour et al., 2012). In all cases, all forces and displacement were concentrated on a  15rather small region of the panel that led to local failure phenomena. All energy was dissipated through only the connections as CLT products are very stiff (Ceccotti et al., 2006). Tests in Ljubljana and San Michele all’Adige demonstrated a number of failure modes in the connection. Failure occurred in the bracket, fasteners, and wood. The most common observed failure in CLT connections is pull-out of fasteners. Pull-out failure is the ideal and desired failure mode, as the most energy can be dissipated by deforming plastic hinges as the fasteners bend back and forth while steadily separating from, but not destroying, the CLT panel (Schneider, 2009). It is most important to not damage the primary structure of the building (i.e., the CLT) as this is much more difficult to repair or replace than the fasteners. Under cyclic loading nails/screws can shear off the head. The most concerning form of failure is failure of the wood panel. If the pull-out resistance is higher than the tensile strength of the CLT panel but lower than the strength of the bracket, group tear-out failure in the CLT panel will occur. It is a non-predictable failure mode and depends also on defects in wood fibers. Group tear-out often destroys the CLT panel and is difficult to fix. Group tear-out failure appears in two ways in CLT structures. The wood fibers within a board can fail, which causes a rupture, brittle block failure. CLT is arranged in several board layers, a second type of group tear-out failure can be observed when the bond between the board layers fails and entire board section gets pulled out. If the pull-out resistance is high enough and the wood remains intact, the weakest point in the connection is the bracket. It will fail at the cross-section with the smallest net steel area. However, it is a ductile failure and the damage rate is not as high as block failure, but does not provide the most resistance to dissipate energy. The bracket failure could also occur in a way that the bolt to the foundation gets pulled through the bracket hole. Depending on the ground motion, the steel bracket sometimes experiences high deformation and plasticization. This is not considered as total failure as long as the bracket does not detach from the panel.  1.7 Damage evaluation  An important part of the observed damage is to find measurements to quantify damage. Over the last years a number of damage indices have been proposed. Williams and  16Sexsmith (1995) discussed various approaches for damage indices. They classified the local damage indices into groups of • Non-cumulative indices • Deformation-based cumulative indices • Energy-based cumulative indices • Combined indices Ductility (defined in terms of rotation, curvature or displacement) and inter-storey drift are widely used to assess structures with non-cumulative indices (Williams et al., 1995). Non-cumulative indices neglect the effect of repeating cycles that occurs in earthquakes. However, based on the simplicity and ease of interpretation non-cumulative indices are widely used (Williams et al., 1995).  Wang et al. (1987) proposed a simple deformation-based cumulative damage model. It was assumed that the development of damage depends on the maximum deformation occurring in a cycle. Further, the rate of accumulation of damage was assumed to be proportional to the damage already occurred.  Gosain et al. (1977) defined energy absorption as a measure of damage. Hysteresis loops, which dropped below 75% of the yield value after reaching the yield value, were negligible for the remaining capacity of the member. Krätzig et al. (1989) developed a more complex energy formulation based on following half-cycles. The first half-cycle of loading at given amplitude is called primary half-cycle (PHC). The subsequent part of the cycle after peak load is called follower half-cycle (FHC). Cumulative formulations for the positive and negative part of the hysteresis loops were calculated separately and then added up to obtain the overall damage index D. To date, most available damage models were developed for reinforced concrete and steel buildings (Liang 2011). Park et al. (1985) proposed a damage index based on existing data from an array of tests conducted around the world. The approach was a combination of both, the maximum deformation and the influence of repeated cyclic loadings. This model is one of the best known and most widely accepted cumulative damage indices for reinforced concrete. Only a few attempts have been made to apply this index to wood-frame structures and cross laminated timber structures (van de Lindt et al., 2006). Van de Lindt (2005) has  17extended the model of Park et al., which was expressed as a function of the sheathing perimeter nail spacing.  1.8 Objectives and scope  The focus of this research is on the connections for timber-steel hybrid structures, and their behaviour under monotonic and cyclic loadings. The research work includes experimental investigation of conventional connections through damage analysis and assessment and comparison between modeled and experimental work. Further, connection methods for such hybrid assemblies are reconsidered and a novel connection method is developed and investigated. To meet these objectives, several milestones were achieved:  a) Experimental Tests: • Performance of experimental monotonic and cyclic tests on conventional bracket-type connections for CLT wall panels used in combination with a steel frame base. This stage includes analysis of test results and determination of elastic-plastic properties. • Assessment and evaluation of failure modes occurring in conventional connections. Includes investigation of influencing parameters on monotonic and cyclic failure. • Analysis of energy dissipating properties within the context of an energy-based damage accumulation model to define a damage index and scale for tested connections. • Comparison of individual connection tests with wall assemblies to study correlation between small-scale tests and full-scale tests. b) Numerical Analysis: • Development of a finite element (FE) numerical model with non-linear springs capturing the behaviour under cyclic loading on the basis of empirical data obtained from the testing of conventional connections. Includes aligning the hysteretic model with a 10-parameter model.  18• Recalculation of the energy-based damage accumulation indices using the FE model results. • Comparison of damage indices from experimental and numerical results. Includes evaluation of the model using the energy-based damage accumulation method and the commonly applied energy equivalent elastic-plastic method. c) Novel Connecting Approach • Analysis of existing connecting methods and evaluation of requirements for an improved connection for timber-steel hybrid infill wall assemblies. • Design of a new connector to meet all defined requirements. Experimental tests to prove relevant performance criteria under monotonic and cyclic loading. • Evaluation of the obtained results in comparison to existing bracket test results.  1.9 Thesis structure  In Chapter 1, different hybridization types are discussed. Structural elements of a building are described with emphasis on connection methods. Finally, there is an overview of methods to determine damage, damage assessment, and evaluation of damage. Chapter 2 describes the conventional connection test series and damage analysis for those tests. Details of the framework of the connection tests are presented. The test results are discussed and further applied to a damage accumulation principle. Comparison of visual test observations and calculated damage indices is presented and a preliminary damage prediction scale is established. In Chapter 3, numerical models of the previously tested and analyzed connections are described. A performance assessment using a damage accumulation principle of the modelled connections is conducted and compared with the tests results. To evaluate the conventional analysis method for connections according to the North American standard, the performance assessment method is compared with the conventional analytical method. In Chapter 4, the development of a novel connecting method for timber-steel hybrid buildings with CLT infill walls is presented. After a short introduction, an overview of ductility in timber structures is presented. To understand the motivation for developing a novel approach for connecting CLT infill walls, conventional connections and the associated  19problems are discussed. It is followed by the description of the new connection design, assembly, materials, and test methods. The results under monotonic and cyclic loading protocols are analyzed and discussed. To demonstrate the benefit of the new design, three variations on the novel connection design are compared to three conventional connections. In Chapter 6, the most essential findings of this research are summarized. The presented overview gives the reader an understanding of connection behaviour under monotonic and seismic loading, damage and damage assessment, and a novel approach for connecting CLT infill walls effectively and efficiently.  202    Chapter:  Damage assessment of connections used in cross laminated timber subject to cyclic loads  Wood structures have generally performed well when subjected to strong earthquakes (van de Lindt 2005; Dujic et al. 2004; Rosowsky et al., 2002). This is achieved through a high strength-to-weight ratio of wood, structural redundancy and energy dissipation capacity (Rainer et al., 1999). The current BC building code for multi-storey wood frame buildings limits the building height to six storeys (BCBC, 2010). The incorporation of a steel frame structure with CLT infill walls can provide a solution to mid-rise and high-rise buildings (Dickof et al. 2012; Stiemer et al. 2012). In order to use CLT in this hybrid structure, it is important to have a comprehensive understanding of its structural properties and connection details (Schneider et al. 2012). In this proposed hybrid application, one leg of the bracket is fastened to the infill panel, while the other side is bolted to the steel frame (Figure 2.1). Bracket types can vary in size, shape, and number of provided holes. Different CLT connections were subjected to monotonic and cyclic test protocols, and a mechanistic damage accumulation assessment model, according to Krätzig’s energy-based index (Williams et al., 1995), was developed for each connection. In this research, the connection details and mechanistic damage index will be calibrated through experimental work. The hybridization utilizes the best engineering properties of each material while improving the pace of construction. However, there is still a challenge to connect wood and steel appropriately due to different thermal expansion coefficients and hygroscopic behaviour of wood. At the interface of infill wood walls with surrounding steel skeleton elements, a temperature drop can cause significant pressure on the wood (Slavid, 2005). Similar conditions with internal forces can be generated when wood swells due to direct contact with water (Slavid, 2005).   21 Figure 2.1    Proposed timber-steel hybrid structure  Traditional timber-frame structures use connections between studs and sheathing that are ductile and provide a variety of load paths (Rainer and Karacabeyli, 1999). Blass and Uibel (2007) considered and tested dowel-type connections in the narrow side as well as the face side of the CLT panel. Based on the empirical data Johansen’s theory (Johansen, 1949) was modified and calibrated with the CLT test results. Energy dissipation in CLT shear walls is achieved through ductile behaviour of connections as CLT products are very stiff in plane (Ceccotti et al., 2006; Popovski et. al. 2010). The most common observed failure mode in CLT connections is fastener yielding combined with pull-out of fasteners from the CLT panel after a large number of reversed cycles and/or some fatigue failure of the fasteners (e.g. Schneider et al. 2012). Yielding of the fasteners is the ideal and desired failure mode, as energy can be dissipated by forming plastic hinges in the fasteners. The most undesirable and unpredictable failure mode is the failure of the wood panel. For each type of physical damage observed, it is of interest to develop a damage index, calibrated with the experimental field observational data (e.g., van de Lindt 2005). Cosenza et al. (2000) and Williams et al. (1995) have discussed various global and local damage indices. Global damage indices describe the overall damage state of a structure.  22Whereas, local damage indices describe the damage which occurs in an individual member or joint (this applies to the connections results provided in this research). The local damage indices can be further categorized (Williams et al., 1995):  i) non-cumulative indices (e.g., Banon et al. 1981);  ii) deformation-based cumulative indices (e.g., Wang and Shah 1987);  iii) energy-based cumulative indices (Gosain et al. 1997; Krätzig et al. 1989), and  iv) combined indices (e.g., Park and Ang 1985). In previously reported studies, damage indices were computed, after the entire loading protocol was completed (e.g. Liang et al., 2011). In this research, the connections are tested to study their mechanical performance under vertical (parallel to the grain) and horizontal (perpendicular to the grain) loadings. The focus is on damage accumulation over time and correlation of a computed damage index and observed connection damage. Furthermore, the data from the experimental work was used to calibrate Krätzig’s energy-based model and a preliminary damage scale was proposed accordingly. The goal of damage indices is to provide a means of quantifying numerically the damage under earthquake loadings (Williams et al., 1995)  2.1 Framework for connection tests and quantification of damage  Six connections comprising combinations of two brackets and five types of fasteners, were tested in the parallel and perpendicular to grain directions of the outer layer of the CLT panels. The test protocol followed ASTM 2126-09 (2009), more specifically the CUREE loading protocol, with displacement controlled monotonic and cyclic loadings. Time, displacement, and force were monitored at each incremental displacement step.  Concurrently, all tests were recorded on video for detailed visual analysis. The test data results were processed and the damage accumulation index at each time step was calculated and plotted versus time. The results were compared to identify significant performance changes in the connections and corresponding damage accumulation index. Based on the comparison, a damage scale with five levels (None, Minor, Moderate, Severe, and Collapse) of damage was defined. The damage scale was generated using twenty-four  23tests were used to generate, and the results of the other 37 cyclic tests were used to validate the generated damage scale.  2.1.1 Connection tests 2.1.1.1 CLT samples  The CLT samples had a dimension of 17.8 cm × 35.6 cm (7" × 14"). The material was three-ply with a layer arrangement of 30-34-30 mm (Figure 2.2). Each layer was oriented with the grain at 90 degrees to the previous layer. The CLT test panels were produced by the cross laminated timber manufacturer KLH in Austria. All boards were from spruce wood, classified S10 according to the European standardization. The moisture content measured was 9.87%, while the average density of the panels was measured to be 408 kg/m3 (ASTM 2395-07a, 2007). Specimens were cut randomly without concern for gaps in the boards out of a 2.3 m × 2.3 m panel. Two different types of samples were considered: the first one has the grain of the outer layers parallel to the longitudinal direction of the panel and the second one has the grain of the outer layers perpendicular to the longitudinal direction of the panel. At the top end of the specimens, four 12.7 mm (½") holes were drilled to attach the CLT sample to the test device.  24 Figure 2.2    Cross section of used CLT (board layers: 30 mm – 34 mm – 30 mm)  2.1.1.2 Brackets and Fasteners  The brackets used for this research are manufactured by SIMPSON Strong Tie, “90 × 48 × 3.0 × 116” (Bracket A, Figure 2.3) and “BMF bracket 105 with stiffening rib” (Bracket B, Figure 2.4). “Bracket A” was connected with one leg to the steel frame using three 12.7 mm (1/2") bolts and the other leg was connected to CLT panel using up to 18 small diameter screws or nails. “Bracket B” was connected with one leg to the steel frame using three 9.5 mm (3/8") bolts and 10 small diameter nails. Both bracket types were cold formed and the steel was 3mm thick. Bracket B comes with a large stiffening rib. Five different fasteners (Figure 2.5) were used in combination with the two brackets. A variety of screws and nails were chosen according to previously reported test series (Ceccotti et al. 2006; Dujic et al. 2004). Table 2.1 summarizes the list of tests undertaken with the two brackets and fasteners.     25  Figure 2.3    Bracket A  (Simpson StrongTie)   Figure 2.4    Bracket B (Simpson StrongTie)    26Image of fastener Description Test initial  a) Spiral nail 4.2 mm × 89 mm  (16d × 3 1/2")  (N)  b) Ring Shank nail 3.76 mm × 76 mm (10d × 3")  (R)  c) Ring Shank nail 4.2 mm × 60 mm (16d × 2 5/16” )  (r)  d) Screw 5 × 90 mm  (S)  e) Screw 4 × 70 mm  (s) Figure 2.5    Applied fasteners with brackets  27Table 2.1    Overview of tested samples and their features Fastener layout (Front view of the bracket) Bracket type Fastener type No. of fasteners Test ID parallel to the grain mono. (cyclic) Test ID perp. to the grain monotonic (cyclic)  A Spiral nail 16d × 3 1/2" 18 A-N-L-M5 A-N-L-M6 A-N-L-M7 (A-N-L-C1) (A-N-L-C2) (A-N-L-C3) (A-N-L-C4) (A-N-L-C5) (A-N-L-C6) (A-N-L-C7) A-N-P-M4 A-N-P-M5 A-N-P-M6 (A-N-P-C2) (A-N-P-C3) (A-N-P-C4) (A-N-P-C5) (A-N-P-C6) (A-N-P-C7)  A Ring shank nail  10d × 3" 12 A-R-L-M0 A-R-L-M5 A-R-L-M7 (A-R-L-C4) (A-R-L-C5) (A-R-L-C6) (A-R-L-C7) (A-R-L-C8) A-R-P-M5 A-R-P-M6 A-R-P-M9 (A-R-P-C1) (A-R-P-C2) (A-R-P-C6) (A-R-P-C7) (A-R-P-C8)  A Ring shank nail 16d × 60mm 12 A-r-L-M4  A-r-L-M6 A-r-L-M9 (A-r-L-C1) (A-r-L-C2) (A-r-L-C3) (A-r-L-C4) (A-r-L-C5) A-r-P-M0 A-r-P-M7 A-r-P-M8 (A-r-P-C3) (A-r-P-C4) (A-r-P-C5) (A-r-P-C6) (A-r-P-C7)  A Self-drilling screw  5 × 90 mm 9 A-S-L-M0 A-S-L-M1 A-S-L-M2 (A-S-L-C7) (A-S-L-C8) (A-S-L-C9) (A-S-L-C10) (A-S-L-C11) A-S-P-M0  A-S-P-M7  A-S-P-M8  (A-S-P-C5) (A-S-P-C6) (A-S-P-C8) (A-S-P-C9) (A-S-P-C10)  A Self-drilling screw  4 × 70 mm 18 A-s-L-M0 A-s-L-M4 A-s-L-M5 A-s-L-M6 (A-s-L-C4) (A-s-L-C5) (A-s-L-C6) (A-s-L-C7) (A-s-L-C8) A-s-P-M5 A-s-P-M6 A-s-P-M9 (A-s-P-C3) (A-s-P-C4) (A-s-P-C7) (A-s-P-C8) (A-s-P-C9)   B Spiral nail 16d × 3 1/2" 10 B-N-L-M4 B-N-L-M5 B-N-L-M6 (B-N-L-C1) (B-N-L-C2) (B-N-L-C3) (B-N-L-C4) (B-N-L-C5) B-N-P-M1 B-N-P-M2 B-N-P-M3 (B-N-P-C2) (B-N-P-C3) (B-N-P-C4) (B-N-P-C5) (B-N-P-C7) A=Bracket A, B=Bracket B, N=spiral nail, R=ring shank nail (long), r=ring shank nail (short), S=screw 5x90mm, s=screw 4x70mm, L=longitudinal to the grain, P=perpendicular to the grain, M=monotonic loading, C=cyclic loading, 1,2,3, ... = test repetition   282.1.1.3 Test Setup  To measure the displacement parallel to the grain and perpendicular to the grain, two different test setups were required. The top end of the CLT specimen was connected to the testing apparatus by friction and shear, using steel clamping plates and 12.7 mm (1/2") bolts as illustrated in Figure 2.6. For the perpendicular-to-the-grain set up, the original settings had to be modified in order to accommodate the rotated bracket (Figure 2.7). Based on the rotated orientation of the bracket, the CLT sample was rotated too and thus, there is only the middle timber layer oriented in the strong direction of the applied force. This creates a much higher bending moment within the CLT specimen. To prevent twisting of the specimen, a roller on the opposite side of the hold-down bracket was installed (Figure 2.7).  Figure 2.6    Test Set-up for connection tests parallel to the grain   29 Figure 2.7    Test Set-up for connection tests perpendicular to the grain with roller in the back to prevent sample from bending  2.1.1.4 Loading Protocol  The connection tests were conducted under monotonic and cyclic loading. For the first set of tests, the monotonic program was carried out with a unidirectional and downwards (tension) loading at a rate of 6.35 mm (1/4") per minute. All tests were displacement-controlled. The testing stopped when a 50% drop from the peak load was reached. The cyclic loading tests applied the CUREE loading protocol, which is a displacement-controlled loading procedure with displacement cycles grouped in phases at incrementally increasing displacement levels (ASTM 2126-09, 2009). Each phase consists of a primary cycle at a certain fraction (percent) of the reference deformation. The reference deformation is given by the value that has not yet dropped below 0.8Ppeak of the monotonic test (Figure 2.8). The consecutive cycles of each phase are on a 75% level of the related primary one. Each subsequent phase consisted of a primary cycle with an increase in amplitude of 20% over the previous primary cycle, followed by two trailing cycles with  30amplitude of 75% of the primary one. If the CLT panel survived at the end of phase 8 (100%), additional phases were added to ensure that it eventually failed. The rate of displacement was chosen at 2.54 mm/s (0.1" per second).  Figure 2.8    Load-displacement curve, equivalent energy elastic-plastic (EEEP) curve of a cyclic test and the reference deformation at 0.8Pmax  The test loading protocols for parallel- and perpendicular-to-the-grain were distinctively different to reflect how the panel would be installed in reality (Figure 2.9 and Figure 2.10). In a construction installation, a wall panel with its outer grain orientation vertical would be held in place with connecting brackets to the steel frame. Thus, vertical movement (i.e., parallel-to-the-grain) would be possible only in an upward direction because the steel frame would prevent downward movement. To accommodate this physical restriction in a testing environment, the cyclic loading protocol parallel-to-the-grain was not allowed to go below zero (Figure 2.9). The loading protocol perpendicular to the grain represents horizontal movement of the wall, which can be positive or negative (Figure 2.10).  31 Figure 2.9    Cyclic displacement schedule CUREE test protocol parallel to grain   Figure 2.10    Cyclic displacement schedule CUREE test protocol perpendicular to grain   322.2 Test results  2.2.1 Hysteretic response  The hysteretic responses of the different connection details, parallel and perpendicular to the grain, are shown in Figure 2.11 and Figure 2.12, respectively. Table 2.2 summarizes the ductility ratios as well as the elastic shear stiffness of tested connections. The ductility ratio was calculated with ∆u and ∆yield, where ∆u is the displacement at an 80% load of the maximum load after reaching maximum load. In order to find ∆yield, the yield load Fyield and the plastic portion of the EEEP curve has to be determined. The plastic portion is the horizontal line equal to Fyield as shown in Figure 2.8. The equation to determine this yield plateau is given in ASTM 2126-09 (2009) with [1]  2 2yield u u eeAF KK = ∆ − ∆ −    where A = the area under the curve from zero to ultimate displacement (∆u). Ke is the elastic shear stiffness which is given as  [2]  max0.4eeFK =∆  . ∆e is the corresponding displacement at 0.4Fmax.  Bracket A in combination with spiral nails (6.20 parallel and 4.83 perpendicular) and 60mm ring shank nails (6.36 parallel and 4.80 perpendicular) showed the highest ductility in parallel- as well as in perpendicular-to-the-grain directions. Ductility and maximum force (Table 2.3) are the governing values to determine the performance. The ranking based on ductility (parallel-to-the-grain) looks like A-r-L-C (6.36 (100%)) > A-N-L-C (6.20 (97.5%)) > B-N-L-C (5.60 (88.1%)) > A-R-L-C (5.44 (85.5%)) > A-S-L-C (3.47 (54.6%)) >A-s-L-C (3.37 (52.9%)). The ductility ranking for perpendicular-to-the-grain is A-N-P-C (4.83 (100%)) > A-r-P-C (4.80 (99.4%)) > B-N-P-C (4.39 (90.9%)) > A-R-P-C (4.16 (86.1%)) > A-S-P-C (3.93 (81.4%)) > A-s-P-C (3.90 (80.7%)). The variation in ductility between different connections parallel-to-the-grain ranged by 47% of the maximum value (A-r-L-C) for parallel-to-the-grain and varied only by 19% perpendicular to the grain.   33The elastic shear stiffness is calculated at 0.4 Fmax as the ratio between force and corresponding displacement. Bracket A with spiral nails shows the greatest elastic shear stiffness parallel to the grain (Ke=8.71 kN/mm), where Bracket B shows the lowest value at 3.86 kN/mm (55% less than A-N-L-C). In perpendicular direction short ring shank nails present the highest value at 5.60 kN/mm. A-N-P-C at 5.68 kN/mm is only 10% below the connection with ring shank nails. Table 2.2    Ductility ratio and elastic shear stiffness of the connections  Ductility ratio, ∆u/∆yield(-) Elastic shear stiffness, Ke(kN/mm) Test series Min Average  Max cov* Min Average  Max cov*  Parallel to the grain B-N-L-C 4.70 5.60 5.96 0.09 2.93 3.86 4.82 0.23 A-R-L-C 4.52 5.44 6.19 0.12 4.86 7.74 10.27 0.31 A-r-L-C 4.19 6.36 8.86 0.37 4.65 7.55 10.21 0.34 A-s-L-C 2.32 3.37 4.05 0.13 2.72 5.51 6.82 0.10 A-S-L-C 3.02 3.47 4.16 0.19 4.66 5.06 5.90 0.29 A-N-L-C 4.52 6.20 7.16 0.17 6.26 8.71 10.61 0.22  Perpendicular to the grain B-N-P-C 3.55 4.39 5.84 0.22 2.13 2.36 2.82 0.12 A-R-P-C 3.62 4.16 4.81 0.12 4.41 4.93 5.67 0.09 A-r-P-C 3.21 4.80 6.54 0.27 4.04 5.60 6.62 0.18 A-s-P-C 3.21 3.90 4.57 0.12 3.54 4.90 6.46 0.15 A-S-P-C 3.12 3.93 4.27 0.15 3.49 4.52 5.31 0.24 A-N-P-C 3.70 4.83 6.12 0.19 4.12 5.04 5.68 0.13  * coefficient of variation   34 Figure 2.11    Hysteretic response of different connection test parallel to the grain  35 Figure 2.12    Hysteretic response of different connection test perpendicular to the grain  36Table 2.3 summarizes the measured maximum forces and the corresponding displacements from the performed tests. The ranking of maximum force (parallel-to-the-grain) is A-N-L-C (51.07 kN (100%)) > A-s-L-C (47.90 kN (94%)) > A-S-L-C (46.01 kN (90%)) > A-R-L-C (42.23 kN (83%) > A-r-L-C (34.92 kN (68%)) > B-N-L-C (26.85 kN (53%), for the perpendicular-to-the-grain directions, the ranking is shown to be A-N-P-C (56.10 kN (100%)) > A-S-P-C (51.00 kN (91%)) > A-s-P-C (49.89 kN (89%)) > A-R-P-C (46.05 kN (82%)) > A-r-P-C (42.72 kN (76%)) > B-N-P-C (27.93 kN (50%). In four out of six combinations, the maximum load parallel- and perpendicular-to-the-grain was within a variation of 10%. Only the combinations A-r and A-s showed a higher variation of 22% and 17%, respectively, whereas the related displacements perpendicular-to-the-grain showed generally higher values at maximum load. Considering ductility and maximum load capacity, combinations A-N-L-C and A-N-P-C behaved the best by providing high load values and high displacements along with great ductility.  Table 2.3    Summary of maximum forces and displacements  Maximum force, Fmax (kN) Maximum displacement, ∆ at Fmax (mm) Test series Min Average  Max Min Average  Max  Parallel to the grain Parallel to the grain B-N-L-C 21.67 26.85 30.65 17.66 25.78 31.29 A-R-L-C 36.87 42.23 45.52 13.13 16.84 22.61 A-r-L-C 33.54 34.92 36.27 9.35 13.23 18.46 A-s-L-C 37.01 47.90 54.37 13.55 15.69 20.36 A-S-L-C 40.31 46.01 52.92 15.89 19.87 26.71 A-N-L-C 44.95 51.07 57.92 13.42 20.65 35.30  Perpendicular to the grain Perpendicular to the grain B-N-P-C 24.05 27.93 34.10 16.55 29.48 58.44 A-R-P-C 40.62 46.05 57.34 14.60 22.99 28.42 A-r-P-C 37.23 42.72 48.76 18.60 20.99 21.92 A-s-P-C 43.28 49.89 55.60 17.97 25.16 36.76 A-S-P-C 44.80 51.00 57.34 19.77 24.64 28.85 A-N-P-C 47.57 56.10 64.56 17.43 29.68 75.43   37 2.2.2 Failure Modes in CLT Connection Tests  Depending on the test direction different failure modes were observed. Failure in tests parallel-to-the-grain was dominated by the pull-out failure of the connectors (Figure 2.13a). Based on the type of fastener, the head of the fastener was often sheared off (Figure 2.13b). This happened most often with the self-drilling screws, which are made of hardened steel compared to the nails, and have higher withdrawal resistance, based on the thread geometry. Another common observation in the tests (parallel-to-the-grain) was that the fasteners connected close to the edge of the sample (to emulate real-world shear wall installation) would cut through the wood fibres and “break out” so they were no longer fully embedded. As such, the row of fasteners close to the edge would provide less resistance compared to a fully embedded fastener (Figure 2.13c). Tests perpendicular-to-grain showed additional failure modes. Wood crushing appeared in all tests through the entire depth of the top layer board (Figure 2.13d). All fasteners were either very loose or completely pulled out after the entire loading protocol was completed. In some cases the wooden boards of the top layer failed in a brittle fracture beginning in the area of the bracket (Figure 2.13e). The brackets rotated around the centre of rigidity and demonstrated high plastic deformation. Even with deformation, Bracket A (90×48×3.0×116) performed well compared to Bracket B (BMF 105), which was dominated by fatigue failure in the edge where both bracket legs come together (Figure 2.13f). Overall, the damage perpendicular-to-grain was considerably more severe than parallel-to-grain. None of the tests caused full bracket failure. Given that it is preferred to have a ductile rather than brittle failure mode to maintain resistance and capacity longer under stress, the most desirable failure mode is pull-out failure of the fasteners, which in that mode can still provide resistance even after reaching maximum load [e.g., combination A-N-L-C is a favorable connection as it showed high ductility (Table 2.2) and high maximum load (Table 2.3)]. Failure associated with fracture of the wood or fastener breakage (e.g., heads sheared off) should be avoided.  38 a) Fastener withdrawal  b) Shear fracture of fastener   c) Edge break out as seen from the  underside of the sample  d) Wood crushing  e) CLT panel delamination  f) Bracket fracture Figure 2.13    Different failure mode of connections (Figures show tests parallel and perpendicular to the grain, therefore some brackets are oriented horizontally)   392.3 Damage accumulation principle  Many models have been developed in the past decades to determine a damage index. A simple and commonly used damage model is the non-cumulative damage index. Damage can be defined in terms of rotation, curvature, or displacement. Non-cumulative damage indices neglect cyclic effects and capture only one damage stage. Gosain et al. (1977) formulated a model to describe damage by using energy absorption. Bannon and Veneziano (1982) proposed a damage model where the plastic energy is normalized with respect to the absorbed energy at the elastic limit (Wong et al., 2001). Park and Ang (1985) used a linear combination of deflection and absorbed energy and cyclic loading. Fajfar and Vidic (1994) proposed a model where maximum displacement and the hysteretic energy are used to determine the damage index.  Gosain (1977) uses an approach where energy absorption is used to define damage as follows [3]  ; / 0.75i ie i yi y yFD F FF∆= ≥∆∑  where De = energy-related damage index, Fi = force in i-th cycle, ∆i = displacement in i-th cycle, Fy = force at yielding, and ∆y= displacement at yielding. Hysteresis loops which drop below 75% of the yielding value after reaching the yielding value were negligible for the remaining capacity of the member. Krätzig et al. (1989) developed a more complex energy formulation, based on following half-cycles (Figure 2.14). The first half-cycle of loading at given amplitude is called primary half-cycle (PHC). The subsequent part of the cycle after peak load is called follower half-cycle (FHC). The FHC essentially captures the stiffness and strength degradation. There are different cumulative formulations for the positive and negative part of the hysteresis. For the positive part of the response, the damage index is defined as: [4]  ,p i if iE EDE E+ +++ ++=+∑ ∑∑  where E+p,i = energy in a PHC, E+i = energy in a FHC, and E+f = energy in a monotonic test to failure. For the negative part of the response (D-), the damage index is calculated using the  40same formula only with the negative parameters inserted (i.e., E-p,i-, E-i, E-f). The overall damage index is defined as: [5]  D D D D D+ − + −= + −  where D+ = damage in positive cycle, D- = damage in negative cycle, D+D- = interaction of D+ and D-. The inclusion of the FHC energy in the numerator as well as in the denominator limits the influence to a lower level compared to the primary term. As such, both deformation and fatigue-type damage are taken into account.  Figure 2.14    Primary (PHC) and follower (FHC) half-cycles (Krätzig et al., 1989)  2.4 Comparison between calculated damage index and visual observation  The Krätzig’s damage index was computed using Eq. [2] based on the hysteretic response curves given in Figure 2.11 and Figure 2.12 , and the results are shown in Figure  412.16 and Figure 2.18. Detailed analysis and summary of the damage index is provided in the following subsections.  2.4.1 Analysis and comparison for tests parallel to the grain direction  The grain orientation has a significant influence on the observed damage and failure modes. This will be reflected in the damage prediction and the accompanied calculated damage indices. All connection tests in parallel to grain direction had no major destruction to the wood panel or connecting bracket. Only small plastic deformations (1 mm to 3 mm) in the corner area of the bracket were found.  The equivalent energy elastic-plastic (EEEP) curve is an ideal elastic-plastic curve circumscribing an area equal to the area enclosed by the envelope curve between the origin, the ultimate displacement (0.8Fmax after reaching Fmax), and the displacement axis (ASTM 2126-09, 2009) (Figure 2.8). Following ASTM 2126-09, total failure is considered at the point when the force drops below 0.8Fmax after reaching Fmax (Figure 2.8). In Figure 2.16 damage indices parallel to the grain were plotted over time. Besides the lower and upper boundary, there is also the averaged damage index shown. The vertical line in all plots shows when 0.8 Fmax after maximum load was reached.  The visual observations of all tests parallel-to-the-grain confirmed the assumption to select 0.8Fmax as complete failure. The first visual impression of the connection at this point can be misleading as the visual damage appears moderate. Closer investigation showed progressive destruction by tearing through the wood fibers of the attached fasteners (Figure 2.15). Hence, complete failure can be considered around D = 0.80 depending on the type fastener (Table 2.4). Only the connection performed with the 5 × 90 mm screws demonstrated a higher average damage index of 0.95 at the defined failure value at 0.8Fmax. The relatively low energy dissipation in the following half cycles compared to the primary half cycle can explain this behaviour (Figure 2.16).  42 Figure 2.15    Tearing through of connectors  Table 2.4    Minimum, Maximum and average values at 0.8Fmax   Calculated damage index  at 0.8Fmax Test series Minimum Average Maximum B-N-L-C 0.72 0.80 0.87 A-R-L-C 0.71 0.78 0.86 A-r-L-C 0.72 0.76 0.81 A-s-L-C 0.74 0.81 0.87 A-S-L-C 0.87 0.95 1.03 A-N-L-C 0.72 0.79 0.87  43 Figure 2.16     Cumulative damage index parallel-to-the-grain   442.4.2 Analysis and comparison for tests perpendicular to grain direction  The second part of the damage analysis deals with the behavior when the force is applied in a ninety degree angle to the grain and connection. That represents the movement in the horizontal direction of the wall panel. The center of inertia is in a different position around this axis which appears with different and stronger failure modes. When the load is applied parallel-to-the-grain, the bracket experiences only a translational movement (Figure 2.17 left). By applying the force perpendicular-to-the-grain, a rotational component of movement is added as the bracket is stiffer in that direction (Figure 2.17 right).   Figure 2.17    Combination parallel-to-the-grain (left), and combination perpendicular-to-the-grain (right)  The plots of Krätzig’s damage index for each group demonstrated narrow variation (Figure 2.18). The curves continue on a constant rate until hitting maximum load (vertical dashed line in Figure 2.18. In 80% of the tests the damage factor at maximum load had reached a value of 0.85 – 0.90 (Table 2.5). However, visual observations showed progressive damage before reaching the maximum load.  45Table 2.5    Minimum, maximum and average values at Fmax   Calculated damage index  at Fmax Test series Minimum Average Maximum B-N-P-C 0.80 0.85 0.87 A-R-P-C 0.89 0.92 0.95 A-r-P-C 0.90 0.93 0.95 A-s-P-C 0.89 0.94 0.99 A-S-P-C 0.86 0.90 0.96 A-N-P-C 0.67 0.81 0.89  In comparison, the damage index plots for parallel- and perpendicular-to-the-grain both demonstrated initial steep slopes in the curves. A significant difference can be seen when the D value exceeds approximately 0.60. In general, the initial resistance of the fasteners in the CLT sample with applied force perpendicular-to-the-grain is significantly higher than parallel-to-the-grain in almost all tests (Figure 2.19). Although this is unexpected given that wood is known to have less capacity when force is applied perpendicular-to-the-grain, this result can be explained by the nature of the bracket and test set-up. For the perpendicular to the grain tests, the L-shaped bracket was loaded in the stronger axis. In this direction, the bracket experiences a transversal as well as rotational movement, and it was observed that in this configuration the bracket was very often pushed into the CLT sample resulting in more resistance, but also heightened destruction of the wood. Group A-N-P-C tests were conducted with spiral nails, which were found to be the most ductile connectors of this test series. The eighteen nails per bracket and ductility of the connector created the difference to the other four fasteners used in this test series. The absorbed energy at maximum loads is bigger than at other connections (Figure 2.19).  46 Figure 2.18    Cumulative damage index perpendicular-to-the-grain  47 Figure 2.19    Comparison between cumulative damage indices parallel- and perpendicular-to-the-grain   48In the progressive stage of the test, wood cracks, shear failure of fastener, and extensive wood crushing occur. Based on the observations, the critical point where complete connection failure happens can be determined on the damage factor of 0.80. It can be seen that at this point, at least 80% of the fasteners have been pulled out, extensive wood crushing had occurred or the fasteners failed in shear where the holding head got sheared off. The remaining holding force had dropped significantly at that stage. The kinks in the curves can be explained by the CUREE loading protocol where the first cycle of each loading level is significant higher than the tail cycles of the previous loading level.  2.5 Preliminary Damage Prediction  In order to use Krätzig’s damage index applicable in predicting and evaluating the damage state of CLT connections, the relationship between the calculated damage index (Figure 2.16 and Figure 2.18) and observed damage (e.g. Figure 2.13 and Table 2.6 and Table 2.7) is established in this section. Table 2.6    Damage description for six CLT connections: loading parallel-to-the-grain Test ID Fmax (kN) ∆Fmax (mm) Damage description A-N-L-C6 51.0 20.6 Bracket experiences big plastic deformation and bends out of plane. Big nail withdrawal.  A-s-L-C7 47.9 15.6 No fastener failed in shear. Edge break out was observed on the bottom of the sample A-S-L-C10 46.0 19.8 One of the 9 screws attached failed in shear after reaching peak load. Although of visible plastic deformation of bracket still in acceptable shape  A-r-L-C4 34.9 13.2 Fasteners have loosened up already at an early stage of the loading protocol. 4 fasteners pulled out completely. A-R-L-C8 42.2 16.8 Small plastic deformation of bracket. All fasteners behave strong and all of them are in place after finishing the cyclic protocol B-N-L-C5 26.8 25.7 No shear failure of fasteners. A visible crack in the rip of the bracket could be found. Connection was still good in shape  49 Table 2.7    Damage description for six CLT connections (loading perpendicular-to-the-grain) Test ID Fmax (kN) ∆Fmax (mm) Damage description A-N-P-C6 48.3 -51.4 30.7 -28.6 CLT panel cracked – wood failure. Intensive wood crushing in the area of bracket. No shear failure of fastener observed. A-s-P-C9 50.3 -61.8 27.8 -22.5 Bracket rotates around center of inertia. The bracket leg attached to the wood got pushed into the CLT which caused additional strength in the negative direction. 5 screws failed in total. Damage to the CLT panel significant. A-S-P-C10 47.3 -54.6 25.4 -23.8 2 screws fail in tension, 5 additional failure of screws were observed over the entire cyclic loading. Strong wood crushing in bracket area.  A-r-P-C6 40.2 -45.2 20.5 -21.4 Top layered board of CLT delaminated at peak load. 3 nails sheared off and all other 9 pulled out completely A-R-P-C7 43.0 -49.0 25.6 -20.2 Even with cracks in top layer very strong behavior. 6 nails sheared off. Very intense wood crushing. B-N-P-C5 25.8 -30.0 30.3 -28.6 Shortly after reaching peak load crack in bracket appeared. No shear failure of nails. Bracket is weaker part of the connection. No wood crushing observed.  Based on the observed damage, five damage limit states, None, Minor, Moderate, Severe and Collapse were determined (Table 2.8). The five damage limit states were defined after careful analyzing all cyclic tests. The corresponding proposed Krätzig’s damage index (D) scales are: D < 0.20 (None), 0.20 ≤ D < 0.35 (Minor), 0.35 ≤ D < 0.65 (Moderate), 0.65 ≤ D < 0.80 (Severe) and D > 0.80 (Failure). The observed damage in relation to Krätzig’s damage index can be seen in Table 2.8. From the observed damages states, it was found that up to a damage index of 0.65 the connections can be repaired and strengthened with additional fasteners and do not have to be replaced. Between 0.65 and 0.80, the damage to the connection and surrounding area is significant and the bracket type connection has to be replaced with a new connection at a different location. The proposed prediction scale applying Krätzig's damage accumulation model is necessarily limited to the connections tested; however, it provides a preliminary approach for predicting the behaviour of bracket type connections in combination with CLT.   50Table 2.8    Relationship between damage index and observed damage of connection test Degree of damage Damage description Krätzig’s damage index None No visible damage observed D < 0.20 Minor Minor pull-out of fasteners (20% of fastener length); light plastic deformation of bracket; minor repairs are required 0.20 ≤ D < 0.35 Moderate Visual permanent deflections of bracket; shear failure of up to 2 fasteners; extensive pull-out of fasteners (50% of fastener length); can be fixed and reactivated as a connection 0.35 ≤ D < 0.65 Severe More than 80% of the fasteners failed (shear and pull-out failure); severe crack in bracket; separation of bracket from CLT panel; requires replacement of bracket in different position at CLT wall to be serviceable again; severe wood crushing in outer layer of CLT 0.65 ≤ D < 0.80 Collapse Total or partial collapse of connection (90%  or more fasteners failed) D > 0.80    00.10.20.30.40.50.60.70.80.910 100 200 300 400 500Damage factor [-]Time [s]A-R-L-C40.8 Fmax Figure 2.20    Relationship between visual damge and damage index  51 3    Chapter: Assessment and comparison of experimental and numerical model studies of Cross-laminated timber connections  Earthquake resistant engineering is a major consideration for structures along the west coast of North America. Especially taller buildings need to be properly designed, in order to provide serviceability or life safety in a seismic event. Generally, wood-frame buildings are known to perform well in earthquakes (Rainer et al., 1999); however, there are limitations under code that prevent the use of wood framing in all desired building designs. For instance, the provincial building code of British Columbia (BC) limits multi-storey wood frame buildings to a maximum height of six storeys (BCBC, 2010). The current code is based on design criteria, which are defined by stresses and member forces calculated from prescribed levels of applied lateral shear force (Ghobarah, 2001). Innovative hybrid techniques, where steel frame structures are incorporated with CLT (CLT) infill walls, offer one possible solution to residential and commercial multi-level buildings to overcome the six-storey height limitation (Dickof et al., 2012). Seismic demand and seismic capacity of a structure are important factors for the design procedure. Timber-frame structures are relatively lightweight structures that obtain their good seismic performance through ductile connections between studs and sheathing, which provide sufficient ductility to the shear wall system through a variety of load paths (Rainer et al., 1999). CLT shear wall panels, however, are relatively rigid bodies with no studs or sheathing; therefore, different methods and connections must provide ductility and energy dissipation. CLT wall-to-floor connections are designed using L-shaped steel brackets, which are nailed or screwed to the CLT wall panel on one side, and bolted to the floor on the other side of the bracket. To apply such bracket connections within a CLT-based hybrid structure (Figure 3.1), comprehensive understanding of their structural performance under reversed cyclic loading is required (e.g. (Dujic et al., 2004; Schneider et al., 2012). Performance-based design is a methodology, where structural design criteria have to achieve a certain level of performance (Ghobarah, 2001). Damage, displacement, or drift, which are easily measurable, can be related to such performance objectives. However, to measure and evaluate damage is a more complex undertaking. Damage is influenced by  52accumulation of structural damage, variation of failure modes of the structural components, and number of cycles before failure occurs (Williams et al., 1995). One possible way to assess and evaluate is the introduction of damage indices (e.g. Schneider et al. 2014).   Figure 3.1    Proposed timber-steel hybrid structure (left), detail of connection (right)  Beyond calculating the maximum capacity prior to failure of a connection, it is also important to characterize and understand the path to failure (e.g., What is damage? How can damage be quantified? Are there intermediate damage stages before collapse? How does a load-displacement curve of a cyclic loading protocol relate to damage progression over time?) Over the years many different damage assessment attempts have been made (Table 3.1). Damage indices are categorized into global and local damage indices. Global damage indices describe the overall damage state of a structure, whereas local damage indices  53describe the damage which occurs in an individual member or joint between adjacent members.  Table 3.1    Categories of damage principles Damage Principle Description References Non-cumulative indices The model neglects the effect of repeating cycles that occur in earthquakes (Williams and Sexsmith, 1995) Deformation-based cumulative indices Models connect damage directly to the displacement or rotation of an element or structure (Wang and Shah, 1987) (Williams and Sexsmith, 1995) Energy-based cumulative indices Models consider the energy absorption in a system or element under cyclic loadings (Gosain et al., 1977) (Krätzig et al., 1989) (Williams and Sexsmith, 1995) Combined cumulative indices Combined models consider displacement and energy absorption in one index (Park and Ang, 1985) (Williams and Sexsmith, 1995) (Lindt and Gupta, 2006)  Damage can be measured in relation to curvature, rotation, energy or displacement.  For most damage principles, a damage index D is calculated. The goal of damage indices is to provide a means of quantifying numerically the damage under earthquake loadings (Williams et al., 1995). The damage index has to be calibrated and should range between zero and one. Zero represents no damage, where one is considered collapsed or destroyed. In previous research, the damage index was computed at one point after the entire loading procedure was completed (e.g., (Liang et al., 2011).) Schneider et al. (2014) investigated six connection types and developed a damage scale for Krätzig’s energy-based damage index (Table 3.1). The proposed preliminary damage scale distinguished five damage limit states: None, Minor, Moderate, Severe and Collapse (Table 2.8). The proposed prediction scale, applying Krätzig's damage accumulation model, is necessarily limited to the connections tested; however, it provides a preliminary approach for predicting the behaviour of bracket type connections in combination with CLT.  54Table 3.2    List of previous research on modelling wood connections Author Model  Comment Polensek and Laursen, 1984 Development of a tri-linear curve to describe the backbone curve of nailed plywood-to-wood connection. Pinching* was considered in the model Dolan, 1989 Development of a hysteretic constitutive law based on exponential curves. A hysteresis loop was divided into four segments, which are defined by different exponential equations with four boundary conditions, respectively. Ceccotti and Vignoli, 1990 Creating a hysteresis model for moment-resisting semi-rigid wood joints. Pinching and stiffness degradation** are included Foliente, 1995 A general hysteresis model containing 13 parameters for single and multiple degree of freedom wood joints based on a modified Bouc-Wen-Baber-Noori model (BWBN) The hysteretic constitutive law can generate a smooth and versatile varying hysteresis shape that accounts for nonlinearity, strength and stiffness degradation, and pinching Chui et al., 1998 Development of a detailed nonlinear finite element model for a single-shear nailed wood joint under reversed cyclic loading The nail was modeled as a beam element that incorporates the effects of large deformation and hysteretic nature. The embedment behavior of wood under the action of a nail was described using Dolan’s four exponential segments (1989) Foschi, 2000 Development of a general model based on mechanical interaction between the connector and the surrounding wood medium.  To achieve pinching behaviour as gaps were formed between the beam and the medium, the connector was modeled as an elastic-plastic beam in a nonlinear medium which acted in only compression He et al. 2001 Modification of the model of Foschi (2000) in a three dimension timber light-frame model  Folz and Filiatrault, 2001 A general and simple hysteresis model for timber structures with ten parameters was developed Developed for sheathing-to-framing connector accounts for Nonlinearity, strength and stiffness degradation, and pinching subjected to general cyclic loading were considered. This model has been incorporated into a program called CASHEW. It was developed for cyclic analysis of wood-framing shear walls Chui and Yantao, 2005 Development of a mathematical model based on the previously developed single-fastener finite element model (1998) It is used to predict the moment-rotation response of the timber connection containing multi-fasteners under general cyclic loading *Pinching is a sudden loss of stiffness. It is caused by loosening and slipping of the connection under repeated cyclic loading along with large deformations. **Progressive loss of stiffness in each loading cycle  For generic application of the connections in hybrid structures, it is necessary to develop a component model of the connection for wall and building modeling. It is desirable to find a simple way to model the main features of the connection subjected to general monotonic and cyclic loading protocols. Compared to CLT connections, traditional wood- 55frame construction is a relatively mature system that can provide an important reference for modeling of CLT connections with the program OpenSees and its subroutine called Seismic Analysis of Wood frame Structures (SAWS). Modeling studies have ranged from simple models based on load-deformation relationship from cyclic tests to highly sophisticated models, including detailed nonlinear elements for each fastener (van de Lindt, 2004). The models development over the course of the last decades are summarized in Table 3.2. Although these mechanics-based models can capture some mechanical features of the joints, the real behavior of a nail joint is rather complicated under different loading situations and it is difficult to accurately capture those in a numerical model (Xu and Dolan, 2009). Meanwhile, the simple model generates the same level of accuracy as the sophisticated model and greater computational effort has to be considered with increasing model complexity (Folz et al., 2001). In the following subsection, experimental work and load-displacement curves reported in Schneider et al. (2014) will be used to model the connections in OpenSees with a CUREE 10 parameter model. The finite element model (FEM) is used to predict the load-displacement response and energy dissipation characteristic of the connections under general monotonic and cyclic loading (Shen et al., 2013). The modeled results will be compared with the experimental results by using two different performance assessment methods. In the first approach, ductility ratio, elastic shear stiffness, and the equivalent energy elastic-plastic (EEEP) curve according to ASTM 2126-09 (2009) will be calculated and compared with the ASTM code provision. In the second method, the energy-based damage indices according to Krätzig’s damage accumulation principle will be applied and compared. The proposed damage scale (Schneider et al., 2014) for Krätzig’s damage accumulation principle will be compared with the calculated damage indices generated from the model.   563.1 Experimental test and analytical models  3.1.1 Test outline  Figure 3.2 gives an overview of the procedure which was used for this study. The test series considers the six connections from Chapter 2. Figure 3.3 summarizes the test combinations including the test ID. Used materials and loading protocolsare described in detail in Section 2.1.1Schematic drawings of the set-ups are shown in Figure 2.6 and Figure 2.7.  Figure 3.2    Flowchart to describe the procedure of damage accumulation assessment  57 Figure 3.3    Overview of tested connection combinations and test identification  583.1.2 SAWS hysteretic model  The SAWS model is derived from the load-deformation relationship based on hysteretic shapes obtained from general monotonic and cyclic tests. The required material properties were derived from the CLT handbook (FPInnovations, 2011) and the report “Standard of performance-rated cross-laminated timber” (ANSI, 2012). The SAWS model in OpenSees is a CUREE-10 parameter model, which can take into account highly nonlinear, stiffness and strength degradation and pinching behavior. It can produce smooth hysteretic loops. The envelope curve is defined by an exponential function curve and a linear line, which was proposed by Foschi (1977) [6]  ( ) ( ) ( )0 1 0 0 0sgn 1 exp / , uF x F R S x S x F x D = + − − ≤   [7]  ( ) ( )2 0sgn sgn ,peak u u FF x F R S x x D D x D= ⋅ + − ⋅ < ≤    [8]  0, FF x D= >   The CUREE-10 parameters model was first described by Bryan Folz (2001). Ten parameters are used to control the hysteretic constitutive law, seen in Figure 3.4. F0 represents intercept strength for the asymptotic line to the envelope curve(F0> FI > 0), Du stands for the displacement at peak load (Du > 0), and DF as the failure displacement. S0 represents the initial stiffness of the hysteretic curve (S0 > 0). The stiffness ratio of the asymptotic line to the envelope curve (0 < R1 < 1.0) is given by R1. R2 describes the stiffness ratio on the descending segment of the envelope curve (R2 < 0). R3 presents the stiffness ratio of the unloading segment off the envelope curve (R3<1). FI indicates intercept strength for the pinching part (FI > 0), and R4 the stiffness ratio of the pinching part of the hysteretic curve (R4 > 0) where the pinching behavior is simplified to assuming a parallelogram. The parameters a (a>0) and b (b>0) control the stiffness degradation and energy degradation, respectively. The degrading stiffness KP is based on previous loading history, as given by [9]  ( )0 0 0/ / aP unK S F S b x= ⋅    where xun is the last unloading displacement off the envelop curve.  59 Figure 3.4    Hysteresis model for CLT connection  3.1.3 Analytical assessment Methods  In this research, two performance assessment methods are applied to compare the modeled results with the test results. The first method uses only the envelope curve of the load-displacement graph to assess the performance and neglects influences of individual cycles. The second method, a damage accumulation method is an approach, where the performance of the connection is assessed under consideration of each loading cycle over time. In the following sections the methods are explained in detail.  3.1.4 Ductility and equivalent energy elastic plastic curve  Ductility and elastic shear stiffness are important numbers to assess and rate connections. Since the load-displacement curve does not provide an exact elastic and plastic section, the American Society of Testing Methods (ASTM) provides a method to translate  60the irregular load-displacement curve into a ideally linear elastic-plastic curve, where ductility ratio, yielding force (Fyield) and related displacement, ultimate force (Fult) and related displacement, as well as elastic shear stiffness can be determined and compared exactly. The linear elastic-plastic curve is an idealized assumption of the connection behavior. The ductility ratio Dratio is calculated as the ratio of the ultimate displacement at 80% load of maximum load after reaching maximum load (∆ult) and the displacement at yielding (∆yield). [10]  ultratioyieldD ∆=∆ In order to find ∆yield, the yield load Fyield and the plastic portion of the equivalent energy elastic-plastic (EEEP) curve has to be determined. The plastic portion is characterized with a horizontal line equal to Fyield. The equation to determine this yield plateau is given with where A = the area under the curve from zero to ultimate displacement (∆ult)  * Fyield cannot be larger than 0.8 Fult Ke being the elastic shear stiffness which is given as  [12]  max0.4eeFK =∆ ∆e = corresponding displacement at 0.4Fmax. 3.1.5 Energy-based damage model  The energy-based damage model follows an approach by Krätzig. In Chapter 2, various approaches for energy-based damage accumulation principles are described and discussed. Krätzig’s principle was chosen to be applied in this chapter as well, in order to be able to compare the obtained results with the previous knowledge. [11]  2 2yield ult ult eeAF KK = ∆ − ∆ −   *  61 3.2 Results and comparison of Test and Model  3.2.1 SAWS hysteretic model results  The monotonic and cyclic experimental results of the six connections are used to check the predictive capability of the SAWS model. Similar to the approach applied by Shen et al. (2013), the parameter estimations for the present model applications are shown in Table 3.3 and  Table 3.4. Table 3.3    Parameter estimation of SAWS model for monotonic connections tests Connection type B-N A-R A-r A-s A-S A-N Direction L P L P L P L P L P L P F0 [kN] 30.43 36.73 34.49 47.94 38.69 36.11 33.9 47.85 28.99 37.81 56.47 46.46 FI [kN] - - - - - - - - - - - - Du [mm] 21 38 15 24 16 18 21 25 18 23 21 37 S0 [kN/mm] 4.442 1.933 7.051 5.103 8.332 7.089 6.479 5.825 6.297 6.028 7.179 6.646 R1 0.001 0.001 0.062 0.010 0.010 0.030 0.131 0.010 0.100 0.070 0.010 0.025 R2 -0.56 -0.38 -0.90 -0.84 -1.06 -0.84 -1.52 -0.93 -1.22 -1.13 -1.12 -0.53 R3 - - - - - - - - - - - - R4 - - - - - - - - - - - - ɑ - - - - - - - - - - - - β - - - - - - - - - - - - A,B=bracket type, L=longitudinal to the grain, P=perpendicular to the grain, more information to the used fasteners can be found in Figure 2.5  62Table 3.4    Parameter estimation of SAWS model for cyclic connections tests Connection type B-N A-R A-r A-s A-S A-N Direction L P L P L P L P L P L P F0 [kN] 24.35 24.95 21.71 27.68 33.72 25.60 32.16 41.47 23.86 23.62 47.70 47.30 FI ]kN] 3.5 4 3 4 3 4 4 5 3 4 4 3 Du [mm] 17 32 14 23 12 21 16 26 16 24 20 24 S0 [kN/mm] 6.683 2.811 11.14 6.210 10.20 8.800 9.824 4.362 6.330 4.705 9.100 5.440 R1 0.010 0.010 0.136 0.114 0.01 0.085 0.135 0.117 0.200 0.215 0.030 0.010 R2 -0.08 -0.24 -0.12 -0.26 -0.08 -0.21 -0.08 -0.25 -0.14 -0.23 -0.13 -0.16 R3 0.95 0.95 1.00 1.00 1.00 0.95 1.00 1.80 1.00 0.95 0.95 0.95 R4 0.01 0.01 0.07 0.01 0.006 0.01 0.008 0.015 0.015 0.017 0.005 0.015 ɑ 0.5 0.5 0.5 0.6 0.5 0.5 0.4 0.6 0.5 0.5 0.45 0.5 β 1.02 1.05 1.02 1.05 1.20 1.05 1.05 1.05 1.02 1.05 1.03 1.05 A,B=bracket type, L=longitudinal to the grain, P=perpendicular to the grain, more information to the used fasteners can be found in Figure 2.5  3.2.2 Hysteretic Response  Cyclic loadings were performed for all test combinations and for load orientation (parallel- and perpendicular-to-the-grain). Besides generating the EEEP curves, ductility, elastic shear stiffness, maximum load and related displacement were analyzed. The EEEP curves, separated by direction, parallel- and perpendicular-to-the-grain are illustrated in Figure 3.5 and Figure 3.6. The summarized values are shown in Table 3.5 and Table 3.7. In the cyclic test results, Bracket A, combined with 18 spiral nails, showed highest ductility values in both directions (parallel-to-the-grain D = 6.2 and perpendicular-to-the-grain D = 4.9). The elastic shear stiffness was calculated at 0.4Fmax as the ratio between force and corresponding displacement. Bracket A with spiral nails showed the largest elastic shear stiffness parallel-to-the-grain (Ke = 8.7 kN/mm), where Bracket B showed the lowest value, at 3.9 kN/mm. The long ring shank nails present the highest elastic shear stiffness value, at 7.8 kN/mm (short ring shank nails Ke = 7.2 kN/mm). The ductility variation between test and model ranges from 2.8% (e.g. A-S-L) to 30.6% (A-N-L). The average variation over all 12 combinations was calculated to 16.8%. The greatest variation can be explained by examining  63the hysteresis curves (Figure 3.5 and Figure 3.6). In some cases the load required to reach the plastic plateau (horizontal part of the curve) varies between test and model up to 12 % on the positive side (tension) and 18% on the negative side (compression) of the graph. The average variation is calculated to 7%. The model lines up with the test result accurately until maximum load. After that point the test results often show irregular drops which can be created by cracks, abrupt fastener failure or wood degradation. Those events are not captured with the model. The model follows an overall linear degradation of a slope of –R2S0. Calculating the yield point, by using the EEEP curve approach, the sudden drops can lead to a larger variation on the ultimate displacement (∆u) between test and model which has a direct influence on the ductility.   64      Figure 3.5    Equivalent energy elastic-plastic curve for test and SAWS model parallel to the grain  65      Figure 3.6    Equivalent energy elastic-plastic curve for test and SAWS model perpendicular to the grain  66Table 3.5    Summary of ductility ratio, elastic shear stiffness, maximum forces, and displacement at maximum force of the connections under cyclic loading (Parallel-to-the-grain) for four test combinations Cyclic loading Ductility ratio Elastic shear stiffness Maximum force Displacement at maximum force Test ID ∆u/∆yield [-]* Variation [%] Ke [kN/mm]* Variation [%] Fmax [kN] Variation [%] dFmax [mm] Variation [%] B-N-L-C1 5.8 4.8 24.5 17.7 B-N-L-C2 5.8 4.8 24.5 17.7 B-N-L-C3 4.7 3.3 21.7 19.4 B-N-L-C4 6.0 3.5 28.7 31.2 B-N-L-C5 5.7 2.9 25.1 31.3 Average B-N-L 5.6 16% 3.9 28% 24.9 3% 23.5 29% SAWS-B-N-L-C03 6.5 5.0 24.2 16.8 A-R-L-C4 5.4 8.9 42.2 14.4 A-R-L-C5 5.2 10.3 45.1 14.5 A-R-L-C6 6.2 9.3 45.2 14.5 A-R-L-C7 5.8 5.4 37.5 20.8 A-R-L-C8 4.5 4.9 40.5 22.6 Average A-R-L 5.4 19% 7.8 6% 42.1 0% 17.4 20% SAWS-A-R-L-C04 4.4 7.3 42.1 13.9 A-r-L-C1 5.2 8.4 35.4 10.6 A-r-L-C2 8.9 10.2 35.4 11.5 A-r-L-C3 5.4 6.8 32.9 11.8 A-r-L-C4 4.4 5.9 35.9 15.7 A-r-L-C5 4.2 4.7 34.4 18.5 Average A-r-L 5.6 0% 7.2 4% 34.8 5% 13.6 12% SAWS-A--r-L-C03 5.6 7.5 32.9 11.9 A-s-L-C4 4.1 6.8 53.9 16.0 A-s-L-C5 3.5 6.4 45.7 15.5 A-s-L-C6 3.4 6.0 55.0 21.3 A-s-L-C7 3.5 5.7 51.9 20.4 A-s-L-C8 2.3 2.7 34.2 19.5 Average A-s-L 3.4 32% 5.5 24% 48.1 4% 18.5 15% SAWS-A--s-L-C04 4.5 6.8 49.8 15.8  67Table 3.6    Summary of ductility ratio, elastic shear stiffness, maximum forces, and displacement at maximum force of the connections under cyclic loading (Parallel-to-the-grain) for two test combinations Cyclic loading Ductility ratio Elastic shear stiffness Maximum force Displacement at maximum force Test ID ∆u/∆yield [-]* Variation [%] Ke [kN/mm]* Variation [%] Fmax [kN] Variation [%] dFmax [mm] Variation [%] A-S-L-C7 3.0 5.9 52.0 17.3 A-S-L-C8 3.3 4.8 42.4 17.8 A-S-L-C9 3.2 4.9 43.6 15.9 A-S-L-C10 3.6 4.7 51.0 26.7 A-S-L-C11 4.2 5.0 40.7 21.7 Average A-S-L 3.5 3% 5.1 8% 45.9 7% 19.9 20% SAWS-A-S-L-C09 3.4 4.7 42.5 15.9 A-N-L-C1 4.5 6.3 51.0 20.9 A-N-L-C2 7.2 10.6 49.5 19.5 A-N-L-C3 6.8 10.1 44.9 16.6 A-N-L-C4 5.7 10.2 50.2 16.1 A-N-L-C5 7.0 9.7 48.0 19.4 A-N-L-C6 5.3 5.9 57.9 35.3 A-N-L-C7 6.9 8.2 44.9 15.2 Average A-N-L 6.2 31% 8.7 21% 49.5 0% 20.4 4% SAWS-A-N-L-C02 4.3 6.5 49.6 19.6  68Table 3.7    Summary of ductility ratio, elastic shear stiffness, maximum forces, and displacement at maximum force of the connections under cyclic loading (Perpendicular-to-the-grain) for four test combinations Cyclic loading Ductility ratio Elastic shear stiffness Maximum force Displacement at maximum force Test ID ∆u/∆yield [-]* [%] Ke [kN/mm]* [%] Fmax [kN] [%] dFmax [mm] [%] B-N-P-C2 3.9 (6.9) 2.2 (4.6) 27.7 (-29.4) 32.9 (-16.6) B-N-P-C3 3.9 (6.3) 2.1 (4.1) 25.0 (-29.4) 31.8 (-18.0) B-N-P-C4 4.8 (4.1) 2.5 (2.4) 23.7 (-26.0) 28.5 (-17.8) B-N-P-C5 3.6 (6.3) 2.2 (4.2) 25.9 (-35.8) 29.1 (-32.8) B-N-P-C7 5.8 (7.3) 2.8 (5.4) 22.0 (-34.5) 29.2 (-32.0) Average B-N-P 4.3 (5.9) 4% (24%) 2.3 (4.1) 0% (44%) 24.9  (-31.0) 1% (19%) 30.3  (-23.4) 6% (36%) SAWS-B-N-P-C02 4.5 (4.5) 2.3 (2.3) 25.1  (-25.1) 32.0  (-32.0) A-R-P-C1 4.1 (5.9) 4.9 (7.7) 43.4 (-40.9) 24.3 (-17.6) A-R-P-C2 4.5 (3.3) 5.7 (5.2) 41.4 (-46.0) 23.1 (-14.9) A-R-P-C6 3.8 (4.1) 4.9 (6.9) 42.2 (-49.7) 24.5 (-14.6) A-R-P-C7 3.6 (2.9) 4.4 (4.8) 47.4 (-56.6) 28.4 (-26.8) A-R-P-C8 4.8 (3.0) 4.8 (3.9) 42.0 (-51.0) 28.2 (-27.2) Average A-R-P 4.1 (3.5) 20% (3%) 4.9 (5.7) 16% (18%) 43.3  (-48.9) 5% (6%) 25.7  (-20.2) 13% (15%) SAWS-A-R-P-C01 3.3 (3.4) 4.1 (4.7) 41.1  (-45.9) 22.4  (-23.2) A-r-P-C3 6.5 (4.7) 6.6 (5.9) 39.8 (-48.9) 18.6 (-20.8) A-r-P-C4 5.6 (3.1) 6.2 (4.2) 37.2 (-40.4) 20.1 (-21.4) A-r-P-C5 4.4 (5.5) 5.9 (7.3) 41.2 (-41.3) 21.5 (-21.9) A-r-P-C6 4.3 (3.3) 5.2 (5.0) 41.4 (-47.8) 20.7 (-21.5) A-r-P-C7 3.2 (3.0) 4.0 (4.1) 41.2 (-47.9) 21.7 (-21.5) Average A-r-P 4.6 (3.7) 11% (19%) 5.6 (5.3) 2% (15%) 40.2  (-45.3) 0% (7%) 20.5 (-21.4) 2% (2%) SAWS-A--r-P-C05 4.1 (4.4) 5.5 (6.1) 40.3  (-42.2) 20.9  (-21.1) A-s-P-C3 3.2 (4.3) 3.5 (8.1) 51.2 (-58.3) 26.1 (-20.2) A-s-P-C4 4.2 (3.2) 4.8 (5.5) 49.7 (-60.5) 24.7 (27.1) A-s-P-C7 3.4 (3.1) 4.0 (5.0) 52.0 (-64.8) 36.8 (-26.8) A-s-P-C8 4.6 (3.8) 6.5 (6.8) 51.3 (-63.1) 28.5 (-18.0) A-s-P-C9 4.1 (3.8) 5.7 (6.3) 47.6 (-62.4) 23.1 (-20.3) Average A-s-P 3.8 (3.6) 26% (19%) 4.9 (6.3) 39% (43%) 50.4  (-61.9) 2% (15%) 27.8  (-22.5) 8% (17%) SAWS-A--s-P-C03 2.8 (2.9) 3.4 (3.6) 49.7  (-52.7) 25.7  (-26.4) *values in brackets are generated with values from the negative branch of the curve   69Table 3.8    Summary of ductility ratio, elastic shear stiffness, maximum forces, and displacement at maximum force of the connections under cyclic loading (Perpendicular-to-the-grain) for two test combinations Cyclic loading Ductility ratio Elastic shear stiffness Maximum force Displacement at maximum force Test ID ∆u/∆yield [-]* [%] Ke [kN/mm]* [%] Fmax [kN] [%] dFmax [mm] [%] A-S-P-C5 3.1 (3.7) 3.5 (5.7) 48.1 (-53.0) 24.5 (-19.5) A-S-P-C6 4.1 (3.7) 4.5 (5.7) 44.9 (-54.2) 22.8 (-23.7) A-S-P-C8 4.3 (3.9) 4.9 (6.1) 48.3 (-52.6) 22.2 (-24.5) A-S-P-C9 4.1 (3.0) 4.4 (4.7) 45.3 (-55.6) 28.9 (-26.5) A-S-P-C10 4.1 (3.2) 5.3 (5.5) 51.2 (-56.8) 28.8 (-24.9) Average A-S-P 3.9 (3.5) 28% (17%) 4.5 (5.6) 24% (36%) 47.6  (-54.4) 2% (11%) 25.4  (-23.8) 6% (1%) SAWS-A-S-P-C05 2.8 (2.9) 3.4 (3.6) 46.7  (-48.3) 23.8  (-24.2) A-N-P-C2 5.2 (4.9) 5.7 (6.7) 43.4 (-46.2) 17.9 (-19.7) A-N-P-C3 6.2 (4.1) 6.2 (7.3) 48.6 (-49.6) 37.3 (-17.3) A-N-P-C4 4.6 (4.6) 4.7 (7.1) 44.8 (-47.0) 23.9 (-19.8) A-N-P-C5 6.1 (4.0) 5.5 (6.0) 52.0 (-54.8) 40.2 (-23.2) A-N-P-C6 4.6 (3.7) 5.2 (6.1) 52.4 (-54.1) 32.4 (-18.0) A-N-P-C7 3.7 (5.6) 4.1 (7.2) 51.1 (-54.6) 32.8 (-17.4) Average A-N-P 4.9 (4.5) 31% (26%) 5.2 (6.7) 17% (34%) 48.7  (-51.0) 12% (6%) 30.7  (-19.2) 24% (26%) SAWS-A-N-P-C04 3.4 (3.8) 3.8 (4.4) 43.0  (-47.9) 23.3  (-24.2) *values in brackets are generated with values from the negative branch of the curve  The elastic shear stiffness is presented by the initial slope between the origin and the yield point. The EEEP curves of tests and the SAWS model for parallel-to-the-grain direction match very well. Only in the cases of A-r-L and A-s-L, the level of the plastic plateau varies by about 10%. The positive branches of the curves perpendicular to the grain show good agreement between test and model. On the negative side it was found that the elastic shear stiffness is considerably higher than on the positive side. The reason can be found in the loading protocol. The next higher loading step is first applied on the positive side. When it then reverses the cycle into the negative side it has to overcome the plasticized bracket and fasteners which results in a higher force at similar displacement rate.  70 Figure 3.7    Hysteretic response of connection tests and SAWS model parallel-to-the-grain  71 Figure 3.8    Hysteretic response of connection tests and SAWS model perpendicular-to-the-grain for two different bracket combinations  72The measured maximum forces and the corresponding displacements from the performed tests and the model are summarized in Table 3.5. In only one out of twelve combinations, the variation of maximum load between test and model is over 10% (B-N-L-C3). The related displacements show a good correlation between test and model. The average variation was calculated to 5.3%. In the case of A-r-L-C, the variation amounts to 12.5%. Overall, combinations A-N-L-C2, A-R-L-C4, A-S-L-C9, A-r-L-C3, and A-r-P-C5 provide the best agreement of the EEEP curve, elastic shear stiffness and ductility ratio. The hysteretic curves obtained from test and model are presented in Figure 3.7 and Figure 3.8. The overall shape in all connection combinations on the positive (tension) side was captured in most cases with a high precision by the model. In Figure 3.8 (perpendicular-to-the-grain), certain variations between test and model can be observed. The reason can be found in the loading protocol. The first loading cycle always starts in tension. In the reversed cycles around maximum load, big plastic deformations have to be overcome. That difference can be seen in the graph. The parameters of the model are the same in tension and compression, which results in a difference as it is expressed in Figure 3.8.  3.3 Performance assessment using the damage accumulation index  By applying Eq. [4] and Eq. [5] to the obtained hysteretic response from testing and modeling, the cumulative damage index was computed at each time step. The comparative results are plotted in Figure 3.9 and Figure 3.10 respectively. The results will be discussed in the following subsections.  3.3.1 Results for connections parallel to the grain direction  Each graph in Figure 3.9 illustrates three curves, one damage accumulation curve generated with test results, the other damage accumulation curve generated with modeling results, and the loading curve of the connection. The plots show that all damage accumulation curves generated with model results are above the test result curves. In the first section of the curve up to D = 0.15 both curves show very little variation. At larger damage values the variation increases. The modeling curves show an overall stronger increase in the  73damage index D. The shape of the modeling curve shows similarity to the equivalent test curve. The major increases created by the next loading step of the CUREE loading can be found in both curves. However, since Krätzig’s principle considers and accumulates damage from earlier stages, smaller variations add up over time of the entire loading process. The corresponding times for test and model when D = 0.8 is reached are summarized in Table 3.9.  The last two columns of the table presents the ratio between time at D = 0.8 and total length of the loading protocol. Reasons for that increase can be found by closer investigation of the hysteretic response. The envelope curve up to maximum load of the model and test match very well. The largest difference in that section was measured to be 12 % at maximum load (B-N-P-C). The difference of the majority is in a 5% range. Past maximum load, the model follows a linear degradation. The original tests results follow only a vague linear degradation. That behavior explains the differences of the greater increase of the damage index starting at around D = 0.5 and becoming continuously greater. Another factor for the variation can be found in the subsequent cycles of the loading procedure. To generate the observed pinching of the connection, the SAWS model rises at repetitive cycles with a slope of 1/R4S0 (Figure 4.4). It is a linear approximation to the smooth curve observed from the test. By comparing those areas of the hysteretic response, a greater variation was identified. This effect causes a greater increase of the damage curve between the linear jumps, which are created by the next primary loading step. The summations of those two parts, which were identified from the hysteretic response, lead to the variation of DModel to DTest. However, the model overestimates the damage D and is therefore on a conservative side at all stages.  Table 3.9    Time of the loading protocol (parallel-to-the-grain) when damage index reached D = 0.8 Test ID Time when reached Damage index 0.8 [s] Ratio between tD=0.8  and tTotal protocol [%]  Test(*) Model(*) Test Model B-N-L-C3 206 (480) 106 (480) 33.3 42.2 A-R-L-C4 180 (366) 79 (366) 27.1 32.1 A-r-L-C3 222 (387) 111 (387) 22.3 36.9 A-s-L-C4 161 (489) 142 (489) 27.3 38.5 A-S-L-C9 91 (400) 91 (400) 24.0 33.5 A-N-L-C2 206 (473) 132 (473) 50.5 49.0  74  Figure 3.9    Cumulative damage index parallel-to-the-grain with loading curve (red)   753.3.2 Results for connections perpendicular to the grain direction  The damage accumulation curves and the applied CUREE loading protocol (red line) for all six connections perpendicular to the grain for the test and model, respectively are presented in Figure 3.10, while how the loading steps are generated in detail is described in Table 3.10. The value to define the amplitude of the initial cycle of each loading level is a percentage of ∆ult. The general characteristics of the damage accumulation curves can be found as well in the test curves as in the model curves. The steep increases represent the first loading cycle of each load level which lines up with the red curve. The followed by a section of the damage curve, which increases with a lower slope representing the sub-sequential cycles of the loading protocol. There are a few factors that influence the characteristics of the damage curve. To explain the variation between the two curves, the hysteretic response has to be analyzed in detail. The energy dissipation of test and model, which is described by the area enclosed by the hysteresis curves (Figure 3.8), does not correlate perfectly at individual loading steps. On the tension side (positive quarter of graphs in Figure 3.8), the hysteresis loops of SAWS model show similar behavior as the loops in the plots parallel-to-the-grain. On the compression side of the graph, SAWS model does not accurately capture the test result. Except for A-r-P-C, the envelope curve of SAWS model is shifted towards left (more deformation). The contribution of bracket deflection towards the entire deflection is not considered adequately in the model. Hence, the calculated damage index of the test result reaches D = 0.8 at an earlier stage. In addition to that, the strength degradation on the compression side does not create a smooth envelope curve. Often the initial cycles of the next loading steps past maximum load are irregular, where the envelope curve of the model follows a linear degradation. Therefore, the model assumes more energy dissipation on the compression side than actually is achieved. The influence of the “initial cycle”-factor is relatively small. Another factor causing the difference between the DModel and DTest was investigated in the behavior of the subsequent cycles of the loading levels. The damage accumulation curves were obtained from the hysteretic responses perpendicular-to-the-grain, which include loading in positive and negative directions (Figure 3.8). In five out of six graphs, DModel is either equivalent to or above DTest in the first section up to D = 0.25. A-r-P- 76C, A-s-P-C, A-S-P-C, and A-N-P-C continue to show good correlation until D = 0.5 is reached. The envelope curve and loading cycles agree very well with the tests, resulting in good agreement between the damage curves. In the case of B-N-P-C, DModel is close to DTest, but stays at all sections below DTest. In the advanced part, after D = 0.25, DModel increases slower than DTest so that the significant value of D = 0.8 is reached at a later point of the loading protocol. Only the connection combination A-N-P-C (Figure 3.10f) shows small variation over the entire curve. The pinching of the connection is created by those two parallel lines. The loading and unloading of the connection follows the same path over the entire protocol. The model parameters are limited to capture the real pinching, which increases the slope in the advanced cycles of the test. The limitation of SAWS model, lead to an assumption that less energy is dissipated over the course of the subsequent cycles, as there is no variation to capture the test results accurately. Hence, the sections between the initial cycles in Figure 3.10 (steep sloped lines) increase slower than the test curve. At the beginning of the test protocol, the influence of the subsequent cycles is insignificant and therefore good agreement between the two curves can be observed. In the advanced stage of the test, the increasing variation between test and model results in a greater difference, especially since Krätzig’s model approach is an accumulation model, where previous events of the hysteresis curve are accumulated. It was found that the biggest difference between test and model of the subsequent cycles can be found at loading levels 5, 6, and 7 of the CUREE protocol (Table 3.10). Even though the damage level of D = 0.8 in the model was reached at a later point of the loading protocol, in connections B-N-P, A-R-P, A-s-P, and A-N-P, DModel reached 0.8 before starting the next loading level. In that perspective, the differences of test and model are in a reasonable range. Table 3.11 summarizes the times of model and test when D = 0.8 was reached and the ratio between time at D=0.8 and total length of the loading protocol. Four out of six connections vary in a range from 1.5% to 9.5%. Connections A-r-P-C3 and A-s-P-C4 show a higher variation of 14.6% and 11.2%. The average difference was calculated to be 8.45%.  77 Figure 3.10    Cumulative damage index perpendicular-to-the-grain with loading curve    78Table 3.10    Amplitude of primary cycles for CUREE-protocol Pattern Step Minimum Number of cycles Amplitude of primary Cycle, %∆ 1 1 6 5 2 2 7 7.5  3 7 10 3 4 4 20  5 4 30 4 6 3 40  7 3 70  8 3 100  9 3 100 + 100αA (120)  10 3 Additional increments of 100α (20%) up to 160  Table 3.11    Time of the loading protocol (perpendicular-to-the-grain) when damage index reached D = 0.8 Test ID Time when reached Damage index 0.8 [s] Ratio between tD=0.8  and tTotal protocol [%]  Test(*) Model(*) Test Model B-N-P-C2 161 (483) 204 (483) 33.3 42.2 A-R-P-C1 133 (489) 157 (489) 27.1 32.1 A-r-P-C5 92 (412) 152 (412) 22.3 36.9 A-s-P-C3 136 (499) 192 (499) 27.3 38.5 A-S-P-C5 115 (480) 213 (480) 24.0 33.5 A-N-P-C4 225 (445) 218 (445) 50.5 49.0 (*) length of the entire loading protocol in brackets  3.4 Comparison of the model to the assessment method  In this chapter the test results of six CLT connections in two load orientation directions with respect to the grain orientation of the face lamina were used to generate a finite element model in OpenSees. To assess the accuracy of the obtained model results in comparison to the test results, two different assessment methods were chosen. The first method assessed ductility, elastic shear stiffness and maximum capacity. This method focuses on the envelope curve and neglects any influence from hysteretic cycles which occur  79at lower amplitude and do not contribute to the envelope curve of the load-displacement curve. The second method is a damage accumulation approach. The damage accumulation model considers each loading cycle, not only the generated envelope curve. In addition to load and displacement, in this approach the relation between load, displacement, and time is tracked. The damage index D is calculated at incremental time steps, which requires knowledge of the load, displacement, and time relation. By using both approaches, both methods showed their strength and weaknesses in assessing the results. The EEEP curve, and herein calculated ductility, elastic shear stiffness and maximum capacity is a straight forward method, where its results can be summarized and compared easily in a table. Test and model showed reasonable good correlation. The observed differences are small and therefore it was interpreted that this model represents a good correlation to the testes connections. However, the damage accumulation method, where the damage index D was calculated in respect to the time, it was found, that the subsequent cycles have a considerable influence on the accuracy of the damage results. The damage accumulation principle showed that in order to reproduce the exact load displacement relation the model needs to be modified to capture the behavior of the subsequent cycles.  804    Chapter: Novel Steel Tube Connection for Hybrid Building Application  4.1 Hybrid Structures  Timber structures are significantly lighter than concrete buildings leading to smaller forces during an earthquake (Moore, 2000). Timber, however, fails in a brittle manner when loaded in tension or shear. That is why structural timber elements are connected with metal fasteners, which allow for energy dissipation and greater displacement without failing. Damaged timber connection can be replaced to revitalize a building after an earthquake.. While wood light-frame construction has been very successful in the North American low-rise residential construction market, the structural use of mass-timber (e.g. CLT) is just recently gaining popularity. One reason for that is that mass timber provides new opportunities to build higher than the height limitation imposed on light-frame timber structures (six storeys in some jurisdictions and four storeys in others) (e.g. NRCC, 2011; BCBC, 2010).  An alternative could be the use of timber-based hybrid structural systems for mid-rise buildings, in particular involving mass-timber. Incorporating timber and steel in a timber-steel hybrid system is one possible way to overcome the height limitation of light-frame wood construction (e.g. Schneider et al. 2014; Dickof et al. 2014).  A hybrid system consisting of steel moment resisting frames with CLT infill walls is illustrated in Figure 4.1 (left). The potential of that system was demonstrated in previous research (Dickof et al., 2014; Tesfamariam et al., 2014), using the connection test results by Schneider et al. (2014) and calibrated Pinching4 hysteretic model by Shen et al. (2013). These studies also identified the importance to provide a stiff connection between infill wall and steel frame while providing enough ductility to fulfill seismic performance requirements.  Schneider et al. (2014) showed in their reversed cyclic connection tests (Figure 4.1, right), that the CLT panel can potentially be damaged during an earthquake and hamper the potential post-earthquake restoration. In this research, a new tube-based connection is introduced that can avoid damage to the CLT panels.  81              Figure 4.1    Timber-steel hybrid system (left) and observed damage of a conventional hold-down connection under reversed cyclic load (right)  4.2 Ductility in Timber Structures  Wood is a great constructing material with a lot of advantages over other materials (i.e. strength to weight ratio, renewable, good insulating properties). However, due to low plasticization ability of timber in bending, shear, and tension, most timber structures are not able to develop full plastic mechanisms at failure (Jorissen et al., 2011). Brittle failure modes can be seen as row shear-out, group tear-out and splitting (Zhou et al., 2013; CWC, 2009). Wooden structures assembled of mass timber products (i.e. CLT) rely on the connections for ductility. Zhou et al. (2013) tested various combinations of bolted connections and found that specimens with large diameter fasteners failed easily in brittle failure. Jorissen et al. (2011) summarized the need for ductile connections in timber structures in four points: 1. To ensure failure occurs with large deformations 2. To allow stress redistribution within a cross-section or among different cross-sections 3. To allow energy dissipation under seismic loading. Energy dissipation reduces the effect of the earthquake on a structure, leading to an overall better behavior.   824. To increase structural robustness. Ductile behavior is a possible way to ensure the structure can accommodate large displacement and rotation demands caused by sudden failure of a single member within the whole structural system.  In order to define what is ductile or brittle in timber connections, several approaches have been discussed. Jorissen et al. (2011) compared six commonly used methods. Key for a simplified applicable method is the conversion of the obtained test load-displacement curve into a linear elastic-plastic curve. The yield load Fy is a significant point herein. Muñoz et al. (2008) pointed out that the calculated yield load can vary up to 80%, depending on the chosen approach. To be able to evaluate and compare connections and assemblies focusing on ductility versus brittle behaviour, Smith et al. (2006) proposed a classification of failure modes (Table 4.1). Based on the ratio between ∆U and ∆Y, Smith et al. classified the failure as brittle, low-ductility, moderate-ductility, and high-ductility. Those values vary on the method chosen to calculate Fy.  Table 4.1    Classification of failure mode according to Smith et al. (2006) Classification Average ductility ratio, Dav Brittle D ≤ 2 Low-ductility 2 ≤ D ≤ 4 Moderate-ductility 4 ≤ D ≤ 6 High-ductility D > 6  Brühl et al. (2012) used the proposed classification by Smith et al. and evaluated various connections ranging from split rings, shear plate connections, or unreinforced dowel connections to punched metal plate fasteners. The majority of the analyzed connectors were in the range of low-ductility to moderate-ductility. One approach to improve ductility can be, using connector which can easily deform without destroying the surrounding wood. Leijten (1998) and Murty et al. 2008 used hollow steel tubes to obtain failure in the tube rather than the wood. Both researches showed, that tubes can easily achieve a high-ductility classification. Araki et al. (2011) considered a very different method to achieve a more ductile behaviour between posts and beams. The steel  83insert plate is manufactured with slotted holes to allow for rotation. The wooden beam is reinforced with metal sleeves to increase the bearing capacity. High strength, pre-tensioned bolts are installed to guarantee a reasonable level of stiffness. The ductility will occur at higher load levels such as seismic loads. Araki showed that his assembly reduces the often seen pinching behaviour.  4.3 Conventional Connections for CLT Walls  With the introduction of CLT to the construction market, designers face the challenge to connect CLT panels adequately. From a seismic resistance standpoint, a panel-based building is a rigid box and the deformation and energy dissipation is lumped in the connections (Latour et al., 2012). To understand the structural behaviour and performance of CLT shear walls, Dujic et al. (2008) studied CLT connections and wall assemblies under cyclic loadings. It was found that the connectors are the critical element for the capacity of the shear walls.  Bracket-type connections are the most common method to create a connection between walls and floors in CLT structures (Gavric et al., 2015). But none of the currently used connection systems takes full advantage of the crosswise stacked layers of CLT as all bracket-type connections are attached on only to one outer face of the panel. The load path is not centred in the panel and the fasteners behave like embedded cantilevered elements (Figure 4.2 left). In case of a rocking of the shear wall, the fasteners cut through the wood fibres and do not benefit from the cross layering. Fasteners with bigger diameter could improve the bending resistance and therefore the cutting through of fibers. However, the design of the bracket limits the diameter of the fastener to a maximum of 5 mm. Proprietary bracket connectors can be used for timber-based hybrid structures. However, tests under monotonic and cyclic loading caused significant damage to the bracket, fasteners and the CLT panels. In Chapter 2 damage was assessed, compared it with a calculated energy-based damage index at each time step, and then calibrated the calculated damage indices to create a five stage damage scale. Of particular interest were the last two damage stages: a damage index higher than 0.65 indicated that while the connection could  84still carry 50% of its maximum load, but it was irreparable; once the damage index reached 0.8, the connection had failed completely. Another connector, applied to prevent uplift of the ends of a wall, are hold-downs. Hold-downs are brackets with one extended leg attached to the wooden wall panel. Hold-downs are good to prevent uplift, but do not provide stiffness under reversed loading cycles and perform very poorly for energy dissipation as they tend to buckle. Latour et al. (2012) approached the problem of the limited energy dissipation with a new bracket design: so-called XL stubs were welded to thick steel L shape brackets with a weaker section in the corner. Two brackets were attached symmetrically to the panel and a significant improvement of load and energy dissipative capacity was observed. However, this set-up is not suitable for an exterior walls of a tall building as the second bracket cannot be installed due to geometrical restrictions.    Figure 4.2    Pull-out failure of nails in bracket-type connection under longitudinal loading (left); deformed tube connection from Leijten (right)   854.4 Steel tube connectors  4.4.1 Existing steel tube connectors  Most fasteners in timber engineering are round cylindrical elements, e.g. nails, bolts and drift pins (Gagnon et al., 2011; CWC, 2009). These fasteners are normally made of steel with a solid cross-section and work very efficiently as shear connections, also in combination with steel brackets (Schneider, 2009). Another approach that allows an increase in energy dissipation of a fastener is to replace the solid element by a hollow element, e.g. tubes. Leijten (1998) developed a hollow tube fastener with diameters of 18 and 35 mm in combination with densified veneer wood. Each tube was inserted into a pre-drilled hole; to achieve a tight fit, the ends of the tube were pressed and expanded over an attached washer, see Figure 4.2 (right). The tube-type connection was tested in tension parallel to the grain, in a four point bending, and in a set-up as knee joint. The connection showed high stiffness and high ductility. By applying tubes instead of pins in a portal frame, the required number of fasteners was reduced from 16 dowels at 16 mm to 4 tubes at 35 mm while increasing the live load capacity by triple. The large plastic range after yielding provided the high capacity of energy dissipation with low impairing of strength. The ductility levels for the joints were 8 and 15 (Leijten, 1998) effectively placing the fastener in the range of high ductility according to the scale proposed by Smith et al. (2006), (see Table 5.1). Another tube type connector was developed by Murty et al. (2008) who used three sizes of small tubes (6.35, 9.52 and 12.7mm) in combination with spruce lumber and Laminated-Strand-Lumber to create a tension connection with insert steel plates. It was found that these small diameter tubes are an effective means of achieving strong and ductile structural connections. The maximum load capacities in spruce were measured at 4.44 kN (6.35 mm tube), 12.0 kN (9.52 mm tube), and 13.8 kN (12.7 mm tube). However, the arrangement with centered link elements caused the wood element to split at failure loads similar to the behaviour of self-drilling pins in combination with insert steel plates (Schneider, 2009). Both, Leijten (1998) and Murty et al. (2008) applied the steel tube together with a link element and loaded it in shear perpendicular to its longitudinal axis along its  86circumference, similar to traditional dowel type fasteners. One of the drawbacks of these types of connections is that with more drift pins applied it becomes more difficult to install all of them. In addition, in a post-earthquake scenario it is very difficult to assess the damage within the connection. Replacement of deformed fasteners in combination with slotted link elements presents a challenging task as it is very difficult to dismantle fasteners with plastic hinges. While previous research demonstrated the potential of tube-type connections, no research has been conducted on the application of steel tubes as connectors for mass-timber panels, in particular for CLT, where insert tubes would benefit from the crosswise stacked layered structure. The objective of the research presented herein is to assess the feasibility of a tube type connector for CLT wall panels, and compare its characteristics such as damage, stiffness, ductility, and possibility of replacement to conventional CLT connectors.  4.4.2 Novel tube connector  In an attempt to avoid significant damage to the wood, tube-type connectors for joining CLT infill panels to steel-frames are hypothesized to meet several key requirements: • Easy to manufacture, install, inspect, and replace; • High initial stiffness, high strength, high ductility and energy dissipation; • Minimal damage to the timber material under monotonic and cyclic loading. To take advantage of the layered structure of CLT, it seems plausible to consider a connector that is inserted into the material. For ease of manufacture and installation, a round insert is preferred over a rectangular shape as it can be manufactured more easily. Tubes are inexpensive and are commercially available in various sizes. The geometry of a tube embedded in wood provides load distribution along the circumference. A steel tube can be designed in such a way that – when loaded past its yield point – it can deform without crushing or splitting of the wood. In this research, a novel approach is chosen taking advantage of the properties of a tube (see Figure 4.3 (left)). The connection assembly consists of a steel tube, a coupler for 12.7mm bolts, a threaded rod (12.7mm) and two nuts (for 12.7mm rods). Three tube diameters are considered: T2: 50.8mm (2"); T3: 76.2 mm (3"); and T4: 101.6 mm (4"). The  87length of the tube is cut according to the width of the CLT panel, herein 99 mm for a three-ply panel, shown in Figure 4.3 (right). The tube-type connector has a 25.4 mm hole on one side. On the opposite side of the hole, the coupler is welded to the inside of the tube with its main axis pointing to the center of the hole.    Figure 4.3    Considered novel tube connectors (left) and CLT panel (right)  4.4.3 Connector assembly  Installation the tube connector into a CLT panel is illustrated in Figure 4.4. A hole with the diameter of the tube has to be drilled into the face side of the wall panel (1). Another hole (shaft hole) has to be drilled in plane, in the center of the wall panel (2). At the edge of the CLT panel at the interface between steel beam and CLT panel where the shaft hole comes out, a notch with about 20 mm in depth has to be cut (3). For installation, the prepared CLT panel has to be dropped in place on top of the steel beam. The tube connector (4) has to be slid into the CLT panel. The threaded rod (5) is insert from the bottom through the steel beam  88with the nut and washer which sits on top of the steel beam (6). Once tube and rod are connected, the bottom nut (7) can be installed to keep the whole connection in place.  Figure 4.4    Installation steps for tube connector  The design of the novel tube connector is not covered by the Canadian timber design standard CSA086 (CWC, 2009) (nor any other standard); therefore, no guidance on loaded end distances is available. CSA086 does provide the minimum end distance for bolts to be greater than 5d (diameter of the fastener) or at least 50 mm, and for split rings (d = 63.5 mm) to be equal or greater than 210 mm (= 3.3d). Accounting for deflection in the tube, the end  89distance was chosen to be 2.5d. This loaded end distance was later proven to be sufficient but could be revised for future research if larger loads are required.  4.4.4 Materials  The CLT samples were cut from a commercially produced product called CrossLamTM SLT3 (Figure 4.3 right) (StructurLam, 2013). The panels had a crosswise stacked layer set-up of 32 - 35 - 32 mm. The CLT panels were produced of SPF (Spruce, Pine, Fir) Grade No.1 and No.2. The samples were cut randomly from a bigger panel of CLT to account for possible gaps in the board layers. Two sizes of samples with the top layered board in the longer direction were cut: 470 × 305 × 99 mm and 560 × 305 × 99 mm. For the tube, T4, the sample length was increased to 560 mm to provide enough space for the holding fixture (dashed line at top end of sample on the left in (Figure 4.5). The large holes identify the tube connection locations, while the four holes at the top end of the samples represent bolts to fix the specimens onto a universal test machine. The symbol d in Figure 4.5 represents the outside tube diameter.  Figure 4.5    Test specimen with tube insert: side view (top) and cross sectional top view (bottom)  90Tubes of different diameters were cut from cold rolled, electric-resistance welded tubes. The steel was grade AISI 1010 carbon steel with yield strength of 413 MPa.  4.4.5 Methods  The test specimens are depicted schematically in Figure 4.5 and one example inside the test frame is shown in Figure 4.6. The CLT panel was tied to the steel beam with a 12.7 mm rod. For the test set-up, a steel beam W12 × 72 (311.2 mm × 305.8 mm (depth × width) was used to replicate the steel frame structure. In the test set-up, the hole to connect the rod with the steel beam was drilled right beside the web and created a slight eccentricity of the load path. However, measurements showed no influence of the eccentricity on the results. The CLT sample was clamped with four bolts at the top end. Since the load was applied centric to the CLT panel, and not eccentric with conventional brackets, additional guides to restrain the sample from out-of-plane movement were not required. The tests were carried out under displacement control at rate of 6.35 mm/min (1/4"/min) for quasi-static monotonic tests (2.54 mm/s (0.1"/min) for cyclic tests) using a calibrated Universal Testing Machine with the displacement applied at the bottom of the steel rod in a vertical direction. The load was recorded through a load cell and the relative displacement between CLT panel and steel rod was measured with a laser device.   91 Figure 4.6    Test specimen in test frame  4.5 Results and discussion  The test series comprised of a total of 26 specimens (11 for quasi-static monotonic tests and 15 cyclic tests). At least three replicates for each series were tested under quasi-static monotonic loading and five specimens are tested under reversed cyclic loading. The quasi-static tests followed a loading regime according to EN-26891 (Standardization, 1991);  92the reversed cyclic tests followed the CUREE loading protocol according to ASTM E2126 – 09 (ASTM, 2009) which is decribed in Section 2.1.1.4. Figure 4.6 shows the test specimen in the test machine.  4.5.1 Results of the static tests  All load-displacement curves from all three tube diameter groups can be seen in Figure 4.4. The results (mean values and coefficients of variation) of the three connector sizes are summarized in Table 4.2. Fy and ∆y describe the calculated yielding load and displacement at yielding load; Fmax and ∆Fmax denote the maximum load and displacement respectively; Fult and ∆Fult represent the ultimate load and displacement at ultimate load; Ke represents the initial elastic shear stiffness, that is determined as the ratio between Fy and ∆y; D and E represent the calculated ductility ratio (ratio of the ultimate displacement to the yield displacement) and the dissipated energy until Fult, respectively. The results showed a direct correlation between tube diameter and maximum load capacity. Tube size, T2, showed the highest load capacity (58.0kN), followed by tube size, T3, (49.9 kN) and tube size, T4, (43.3 kN). The biggest tube (T4) exhibited the lowest stiffness at 5.6 kN/mm, whereas tubes T3 (16.4 kN/mm) and T2 (18.6 kN/mm) had a significantly higher stiffness. The displacements at maximum load were comparable with 18.3, 25.0 and 17.8 mm for tube sizes T2, T3 and T4, respectively. According to the ductility classification proposed by Smith et al. (2006), T2 and T3 are highly ductile, whereas T4 exhibits only low ductility. A significant plastic plateau occurred at least at half Fmax where the load was constant before increasing to the maximum load, followed by a load decrease and failure (see Figure 4.10).   93Table 4.2    Test results of tube-type connections under static loading  T2-1-M T3-1-M T4-1-M  xmean COV xmean COV xmean COV Ke [kN/mm] 18.6 0.42 16.4 0.54 5.6 0.34 Fy [kN] 52.7 0.07 41.7 0.04 34.0 0.09 ∆y [mm] 3.2 0.39 2.9 0.46 6.3 0.27 Fmax [kN] 58.0 0.09 49.9 0.05 43.3 0.08 ∆F_max [mm] 18.3 0.42 25.0 0.11 17.8 0.07 Fult [kN] 46.4 0.09 39.9 0.05 34.6 0.08 ∆Fult [mm] 22.3 0.54 35.5 0.11 19.7 0.08 D [-] 6.9 0.30 14.5 0.70 3.2 0.22 E [kNm] 1121.7 0.62 1273.4 0.18 1826.2 0.70  All three tubes showed a similar load-displacement relationship, with high initial stiffness, a plastic plateau before reaching maximum load, and a decreasing section before failure. The T3 tube, however, exhibited the most desirable behaviour, as the load maintained at least 35 kN (70% of Fmax) over an extended period, and demonstrated a slow decrease after maximum load. The connection failed in a ductile manner through massive deformation and buckling of the tube before the weld connected to the tube failed. This failure mode occurred in the same fashion for all tube diameters (Figure 4.7), and was depicted in the load displacement curves as a sudden drop in load capacity. None of the tubes caused any form of wood crushing or cracking of the CLT sample (Figure 4.8). The averaged load-displacement curves of all three tube connectors is shown in Figure 4.9. The curve is divided in four stages, which are colored green, yellow, blue, and red.  • Stage I  A rapid increase in load in the linear elastic range of the tube. • Stage II The steel tube starts yielding and the load increase flattens. • Stage III Strain hardening of the steel tube and building up load towards the buckling load. • Stage IV After reaching the buckling load, the part where the coupler is welded onto the tube is pulling down, creating the shape of a funnel until the weld fatigues and starts to break from the outside.  94   Figure 4.7    Tested 2 inch tube connector beside CLT panel (left); Tested tubes 2 inch, 3 inch, and 4 inch (right)   Figure 4.8    Failure mode of a 76.2 mm tube connection under monotonic loading   95  Figure 4.9    Schematic graph explaining the relation between measured load and observed displacement       964.5.2 Results Cyclic Tests  After completing the monotonic test series, each tube size was tested with five replicates with a cyclic CUREE loading protocol that was generated under consideration of the monotonic results. Representative load-displacement curves of each size are presented in Figure 4.10. The envelope curves of all three tube diameters followed very closely the curves from the monotonic curves. However, the maximum loads from previous performed monotonic tests were not achieved with the cyclic loading protocol.  None of the CLT test samples experienced any crack or damage within the cyclic loading procedures, see Figure 4.7. Failure occurred in all tests in a very similar manner. First the tube deformed repeatedly until the tube cracked at the corner between weld and tube.That crack in the steel tube subsequently propagated in the next cycles around, outside the weld until failure. Regardless of the diameter, failure was always found at the edge of the welded coupler and tube. Initial cracking and crack propagation was visually very obvious and is also illustrated in the load-displacement graphs. The weakest point was determined in the tube and not in the weld. Therefore, the weld was not governing the connector strength.  97 Figure 4.10    Monotonic and cyclic load-deformation curve for 50.8 mm tube; 76.2 mm tube; 101.6 mm tube  98The results of tube connections tested under reversed cyclic loading (mean values and coefficients of variation) are presented in Table 4.3. The energy dissipation was measured as a ratio between dissipated energy in cycle i (Ed,i) and the available potential energy (Ep,i). That value represents the equivalent viscous damping ratio νeq. νeq expresses the hysteresis damping property of the connection and is determined at the 1st cycle of each load level. νeq is calculated as the ratio of Ed,i to Ep,i, multiplied by 4π (Chopra, 2006): [13]  ,,,4d ieq ip iEEνpi=⋅ The available potential engery Ep,i can be determined as Ep,i = ½ Fi · ∆i, where Fi is the maximum force in the considered cycle i and ∆i is the related displacement. Since the loading cycle does not have a negative section, νeq,i is calculated by using 4  instead of 2π. All tubes demonstrated very high compressive capacities (revered part of the cycle), which ranged from 90 per cent to 95 per cent of the values in tension (initial loading of the connection). The elastic shear stiffness was the greatest for T2 (23.8 kN/mm), followed by T3 (16.8 kN/mm) and T4 (6.7 kN/mm). The shear stiffness of T2 was found nearly four times higher than T4, whereas only 1.5 higher than T3. Tube sizes T2 and T3 provided a high level of ductility (DT2= 9.8, DT3= 8.8), while size T4 exhibited low ductility (DT4= 3.6). Tube T3 demonstrated the highest viscous damping ratio at 17.6%, followed by T4 (10.6%) and tube T2 (10.2%). ∆Fmax increased with increasing diameter, whereas Fmax decreased at the same time. The elastic shear stiffness decreased linear by with increasing diameter of the tube. The calculated COVs vary up to 16.4%, which is not only influenced by the tube (pure steel failure COVs should be smaller) but also the variation of the weld size. As the failure has happened always on the circumference of the weld, variations in weld size influence the COV significantly.  99 Table 4.3    Test results of tube-type connections under cyclic loading  T2-1-C T3-1-C T4-1-C   xmean COV xmean COV xmean COV Ke [kN/mm] 23.8 0.72 16.8 0.67 6.7 0.40 Fy [kN] 53.1 0.08 34.6 0.21 31.5 0.07 ∆y [mm] 2.6 0.56 2.4 0.61 4.9 0.36 Fmax [kN] 55.7 (-52.1)* 0.08 37.0 (-35.6)* 0.19 36.4  (-33.7)* 0.06 ∆F_max [mm] 13.1 0.42 14.5 0.42 15.7 0.22 Fult [kN] 44.6 0.08 29.3 0.22 28.6 0.07 ∆F_ult [mm] 17.1 0.24 18.6 0.23 17.0 0.18 D [-] 9.8 0.61 8.8 0.54 3.6 0.39 E [kNm] 841.6 0.30 609.8 0.40 458.8 0.26 νeq [%] 10.2 0.24 17.6 0.39 10.6 0.53  4.6 Comparison to conventional connections  For a better evaluation of the performance of the novel tube-type connector, the test results are compared to previously tested bracket connections (Chapter 2).  4.6.1 Failure modes  One main focus of the development of the novel tube connection was to avoid the uncontrolled destruction when bracket connections fail. Schneider et al. (2014) identified six failure modes: Fastener withdrawal, shear fracture of fastener, edge break out of CLT panel, wood crushing, CLT delaminating, and bracket fracture and showed, that those failure modes occurred randomly, regardless of which bracket or fastener was used. Figure 4.2 (left) illustrates the fastener withdrawal failure mode. All tube-type connections, in contrast, failed at the circumference of the welded coupler (Figure 4.7). There was no other failure mode identified and the CLT panel remained undamaged.  100 4.6.2 Static load-displacement curves  Figure 4.11 illustrates the different load-displacement curves for tube connections and three bracket connections. Bracket connections showed already plastic deformation at an early loading stage with irreversible deformations created through pullout of the fasteners or deformation of the bracket. In contrast, tube connections demonstrated linear elastic behaviour in the first stage of the graph. All three tube connections presented a rapidly load increase and reach at least 28 kN within a deflection of only 5 mm. The maximum loads were found at a displacement of 18.3 mm (T2), 25.0 (T3), and 17.8 mm (T4), whereas the bracket connections reached their maximum load generally at a larger displacement (∆A-N = 42.2mm, ∆A-S = 31.2 mm, and ∆B-N = 23.6 mm). The geometry of the tubes, in particular the diameter, influences that the load drops fast after reaching the maximum load. After the tube is buckled, the weld fails which then causes the rapid load drop.   Figure 4.11    Load-displacement relation of static testing of tube-type and bracket type connections  1014.6.3 Hysteretic curves  Figure 4.12 depicts the hysterics curves from three bracket type connectors and three tube-type connectors. As described before, the envelope curve of the hysteretic curves of the tube-type connections followed very closely the load-displacement curve under static loading. In contrast, a larger variation between static and cyclic curves was observed with bracket-type connections. Especially Bracket A with spiral nails stands out with a large variation in the initial slope, which is captured by the elastic shear stiffness. In five out of six connection combinations (except A-S), the achieved maximum load from the static test was not reached. A major difference between tube and bracket connections was observed in the unloading part of the cycle. Connections subjected to cyclic loadings usually show pinching behaviour. Pinching is a phenomenon where the curve does not approach zero on the load axis and deflection axis at the same time while unloading. That occurs because fastener and bracket are plastically deformed and therefore hold up with some force against the reversed cycle. The considered bracket connections showed relatively small pinching. Tube-type connections behave very different from brackets: all three tube connections stand out with their capability of dissipating significant more energy with bigger load amplitudes within the loading cycle.   102 Figure 4.12    Hysteretic curves of tube-type connections and bracket-type connections  103 4.6.4 Equivalent energy elastic-plastic curves  The equivalent energy elastic-plastic (EEEP) curves of the tube connectors compared to those of the three bracket type connectors can be seen in Figure 4.14. All three tubes T2, T3, and T4 demonstrate their high elastic stiffness (Ke). Two of three tubes are higher than all compared bracket connections. However, only T3 reached a big ultimate displacement. T2 and T4 could not reach comparable results to the brackets. The ultimate displacement for connections under cyclic loadingwas found to be slightly smaller compared to their static results. All three tested tube sizes demonstrated high values for Ke under cyclic loadings. The ultimate displacement was relatively limited.    Figure 4.13    Equivalent energy elastic-plastic curves under monotonic loading   104  Figure 4.14    Equivalent energy elastic-plastic curves under cyclic loading  4.6.5 Stiffness and Capacity  Figure 4.15 and Figure 4.16 summarize the test results of all six considered connections. The column diagrams illustrate very well the relationship between tube diameter and maximum load capacity. Compared to bracket connections, the three tubes reached higher maximum loads and exhibited much higher elastic stiffness. It seems, that there is a direct correlation between tube diameter and load capacity. To support this finding more tests need to be executed. In case of the bracket connections, the number and type of fastener influenced significantly the capacity. The connection with the highest load capacity was fastened with 18 spiral nails, whereas the screwed connection in combination with bracket A had only 9 screws. For this research approach the correlation between number of fastener and maximum capacity was not studied. The elastic stiffness (T3) showed a nearly three times bigger value than the highest value of the brackets (A-N). The ductility ratio of T2 and T3 are 6.9 and 14.5, compared to 4.7 (B-N) and 3.8 (A-N).   105The linearity between tube diameter and capacity was not found under cyclic loading. T3 had a disproportional smaller capacity than T2 and T4 under cyclic loading. T2 showed a higher elastic shear stiffness, which turns that parameter closer to a linear relationship between T2, T3, and T4. In the comparison with the brackets, tubes provide in parts (T2 and T3) significant higher elastic shear stiffness. T2 (Ke=23.83 kN/mm) stands out with a nearly three times bigger value than A-N (Ke=8.79 kN/mm).  Figure 4.15    Overview of elastic shear stiffness Ke, yielding load Fy, maximum load Fmax,  ultimate load Fu, and ductility ratio D under static loading   106 Figure 4.16    Overview of elastic shear stiffness Ke, yielding load Fy, maximum load Fmax,  ultimate load Fu, ductility ratio D, and equivalent viscous damping ratio γRd under cyclic loading   1075    Chapter: Conclusion 5.1 Main contributions  Population densification in urban areas is increasingly desired around the world. To promote this phenomenon, there has been growing emphasis on refining the design of tall residential buildings. The design of buildings changes with location and size given the considerably intensified impacts of such factors as snow, wind, and seismic loads as the height of a structure increases.  To broaden future possibilities for innovative and sustainable building methods, wood and engineered wood products are gaining significant popularity as structural building materials (Green et al., 2012). As with any material, however, wood has specific properties that limit its application. To overcome the individual limitations of wood, hybridization with other materials provides a possible solution. Over the years, research has been conducted to explore methods that capitalize on the benefits of hybridization (i.e. Liu et al., 2008; Sakamoto et al., 2004). Regardless of the level of hybridization (through material, system, or structure), the key to success for optimizing the strength, ductility, or other similar properties, is through optimizing the connection between the materials (i.e. Dickof et al., 2012; Moore, 2000). This research focused on the connections of a hybridization method for residential buildings, where the gravity loads are taken by the steel post and beam structure and the lateral stiffness is achieved by solid wood infill walls and floors. Over the past few years, research has been conducted on proprietary CLT connections, which are mainly bracket-type connections. In general, past research focused on the overall behaviour of those brackets until they were considered damaged; however, these investigations relied upon non-standardized qualitative assessments of damage. One of the focuses of this research was to assess and quantify damage for bracket-type connections in CLT panels under monotonic and cyclic loading. Over 100 bracket connection tests were conducted with various fasteners and brackets, in directions parallel to the grain and perpendicular to the grain, and under monotonic and cyclic loading protocols. An energy-based formulation according to Krätzig was applied to calculate the development  108of the damage factor over time, and the resulting factor was characterized and validated with visual observation.  Analysis of the test results showed that five of six types of connections experienced failure when the average damage index D reached approximately 0.80; therefore, for the damage prediction model, complete failure of the connection is assumed to occur at this level. Based on the observed damage and D values, five damage limit states were categorized: None, Minor, Moderate, Severe and Collapse. Upon completion of the bracket tests, damage assessment, and derivation of a damage scale, six of the connections were modeled in OpenSees. For the modeling, a CUREE-10 parameter model was chosen to reproduce the test curves. The obtained load-displacement results from both test and model were analyzed according to the ASTM standard as well as an energy-based damage accumulation principle. The two analyzing methods were used to assess and compare the results from testing and modeling to get a better understanding of the precision of the model. According to the ASTM analyzing method, the overall modeling result correlates with the test results very well. In the ASTM method, ductility ratio and elastic shear stiffness of the envelope curve of the hysteretic results are considered and plotted in an equivalent energy elastic-plastic curve (EEEP). For the energy-based damage accumulation principle, the load displacement results were processed to calculate a damage index at each incremental time step. The applied energy-based damage accumulation method showed a greater variation between test and model than the EEEP curves. It was observed, that the damage index curve of the model in parallel to the grain direction was increasingly stronger for all six combinations. Throughout all six combinations and both loading directions (parallel and perpendicular to the grain) a major difference was found in the damage index development generated by the subsequent cycles, the influence of which are neglected if one considered only the envelope curve of the hysteretic response. It was shown that the influence of subsequent cycles is significant depending on the method of assessing the results. In combinations perpendicular to the grain, the overall curve resulted in an overestimation of the performance of the connection. Therefore, the EEEP curve roughly approximates the performance but analysis of the subsequent cycles is required to better reflect the empirical performance of the connections.  109To this point, this research focused on conventional connection systems (i.e., commercially available brackets and fasteners). When CLT was invented, regular connecting methods from existing techniques were reapplied to the new material. Only limited attempts have been done to date, to improve the connecting methods and in particular take advantage of the structure of CLT with its crosswise stacked board orientation. Through the series of bracket tests, a set of goals was defined to inform the design of a novel connection:  • Easy to manufacture, install, inspect, and replace; • High initial stiffness, high strength, high ductility and energy dissipation; • Suitable for construction in seismically active areas. To address those requirements, a tube-type connection for connecting timber infill walls to steel-framed superstructures was developed. A total of 27 connection assemblies were tested under quasi-static monotonic and reversed cyclic loads. All goals were achieved.  It was demonstrated that this connection has great initial stiffness while providing the required ductility. There was no wood failure observed. All test series failed through ductile steel yielding. By avoiding wood failure, the results had very little variation, which helps for an accurate damage prediction. For the bracket-type connections over 100 tests have been conducted and analyzed. It should be noted, however, that the damage index developed in this research is not yet comprehensive enough to generalize the conclusion, and further tests are required. As well, the damage observations and damage descriptions provided in this research are subject to uncertainty and subjectivity, and should be verified/validated by multiple experts. The comparison between individual connector and CLT walls was based on a limited availability of wall results. The calculated damage curves for the CLT walls can be compared in a clear way with other calculated damage accumulation curves. However, to support the comparison explored in this research the CLT panels must have the same size as well as the same loading protocol. As such, additional testing with the desired conditions would be needed to develop a substantiated general formulation. The application of an energy-based damage assessment method to conventional bracket connections for CLT infill walls is a further area of potential research . Through this type of research, the damage evaluation for individual connections can be applied to  overall  110performance based design and evaluation of hybrid structures. This level of testing for walls and structures would necessarily explore the influence of size, vertical loading, and number of connectors on the damage index. In addition, the comparison between the widely used EEEP method and the energy-based method demonstrated the limitations of the traditional EEEP method to analyze connections. The energy-based method demonstrated particular strength for evaluating the precision of a finite element model. The novel tube-type connection demonstrates great potential as an alternative to the commonly used brackets. It is anticipated that further parametric testing would reveal that by changing parameters such as tube diameter, wall thickness, and grade of steel, such a tube type connection can be adjusted and tailored to meet the requirements of emerging timber-steel hybrid structures especially enhancing performance under seismic conditions. The research would also need to be expanded to investigate overall behaviour of the connection through shear wall tests. With further research, such tube-type connection methods may be incorporated into guidelines such as the technical guide for design and construction of tall wood buildings in Canada, which provides a methodology incorporating new connections in tall wood and potentially hybrid buildings.  This research presented the development of a novel steel tube connector for hybrid wood-based structures. The design goal for the novel tube connector was to provide a connection with high initial elastic stiffness and high ductility. It should be easy accessible for installation as well as for inspection, and - if needed – replacement. Further, the connection should behave as a fuse to dissipate energy and a weak link to fail without damaging the CLT panel. A total of 27 connection assemblies were tested under quasi-static monotonic and reversed cyclic loads. The results can be summarized as follows: • The new tube-type connector demonstrated capability for an easy on-site assembly, as well as easy replacement. Steel frame and CLT infill panels can be prefabricated and assembled fast on site. All three tested diameters showed great linear-elastic initial stiffness, which provides ability to resist smaller seismic impacts without any plastic deformations. Sizes T2 and T3 were classified as highly ductile, demonstrating that  111the connecting system can provide high ductility when appropriate geometric parameters are chosen. The same ductility was achieved under reversed cyclic loading. • The envelope of the cyclic results followed closely the obtained curves from the monotonic tests. Tube-type connections as they have been designed and tested in this study can dissipate a significant amount of energy on the reverse cycle as the threaded rod is stiff enough to deform the deform the tube back to its initial shape.  • All test specimens failed in a ductile steel yielding mode; in no specimens was any wood failure observed. By avoiding wood failure, the results had very little variation, which allows to accurately predicting the type of damage.  In summary, the tube connections showed two major differences when compared to traditional bracket connections: i) the completely linear elastic behaviour at the beginning, and ii) the continued load increase after yielding. Both phenomena are founded in the geometry of that connector effectively making the novel connector a very promising alternative. With further research, such tube-type connection methods may be incorporated into design guidelines for hybrid buildings.  5.2 Recommendations for further research  In the first part of this research, conventional CLT connections were tested and analyzed for possible application in timber-steel hybrid structures. The focus was on damage characterization and prediction, and energy dissipation of the connection. Building on the outcomes of these investigations, several avenues for further research have been identified:  • Applying other damage accumulation principles to the generated data sets and performing a comparative study to confirm/improve the generated damage scale; • Studying and testing other engineered wood products, such as laminated veneer lumber (LVL) and laminated strand lumber (LSL) for possible infill wall panels in combination with conventional brackets as used in this research; • Applying the obtained results to a full-scale timber-steel hybrid infill wall set-up; and  112• Generating a finite element (FE) model and calibrating that model to the full-scale test results. In a second part of this work, a novel tube-based connection approach for fastening infill wall panels to a steel frame was developed and tested. 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Each test provides three graphs. The first one on top shows the load-displacement relationship, the second one below shows the relation displacement over time, and the third one presents the load over time. The labeling of the tests was described in detail in Chapter 2, Figure 2.5 and Table 2.1.                       122Appendix A  Bracket B with 10 Spiral Nails 4.2 mm x 89 mm A.1 Monotonic Test Results   Test ID   B-N-L-M4 B-N-L-M5 B-N-L-M6 Average Fmax [kN] 28.7 32.8 26.9 29.1 dFmax [mm] 22.0 27.0 19.0 21.0 Fult [kN] 23.0 26.3 21.5 23.2 ∆ult [mm] 35.0 37.0 28.0 33.9 D  [-]       4.5 Ke [kN/mm]       3.5 Fy - CUREE [kN]       26.2 ∆y - CUREE [mm]       7.5 Figure  A.1.1    Monotonic test results for connection combination B-N-L-M  123    Test ID   B-N-P-M1 B-N-P-M2 B-N-P-M3 Average Fmax [kN] 28.9 30.7 35.2 31.3 dFmax [mm] 42.0 38.0 39.0 38.0 Fult [kN] 23.1 24.6 28.1 25.1 ∆ult [mm] 55.0 59.0 53.0 54.1 D  [-]    2.9 Ke [kN/mm]    1.5 Fy - CUREE [kN]    28.3 ∆y - CUREE [mm]    18.6 Figure  A.1.2    Monotonic test results for connection combination B-N-P-M  124A.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 24.50 kN (-) - kN dFmax  (+) 17.66 mm (-) - mm 0.1 Fmax (+) 2.45 kN (-) - kN 0.4 Fmax (+) 9.80 kN (-) - kN 0.8 Fmax (+) 19.60 kN (-) - kN D (+) 5.84 [-] (-) - [-] Ke (+) 4.82 kN/mm (-) - kN/mm Fy  CUREE (+) 22.70 kN (-) - kN ∆y  CUREE (+) 4.71 mm (-) - mm Figure  A.2.1    Cyclic test result for test combination CLT-B-N-L-C02 125    MECHANICAL PROPERTIES Fmax (+) 21.73 kN (-) - kN dFmax  (+) 19.42 mm (-) - mm 0.1 Fmax (+) 2.17 kN (-) - kN 0.4 Fmax (+) 8.69 kN (-) - kN 0.8 Fmax (+) 17.39 kN (-) - kN D (+) 4.91 [-] (-) - [-] Ke (+) 3.31 kN/mm (-) - kN/mm Fy  CUREE (+) 19.40 kN (-) - kN ∆y  CUREE (+) 5.86 mm (-) - mm Figure  A.2.2    Cyclic test result for test combination CLT-B-N-L-C03 126    MECHANICAL PROPERTIES Fmax (+) 28.71 kN (-) - kN dFmax  (+) 31.20 mm (-) - mm 0.1 Fmax (+) 2.87 kN (-) - kN 0.4 Fmax (+) 11.49 kN (-) - kN 0.8 Fmax (+) 22.97 kN (-) - kN D (+) 5.96 [-] (-) - [-] Ke (+) 3.49 kN/mm (-) - kN/mm Fy  CUREE (+) 24.13 kN (-) - kN ∆y  CUREE (+) 6.91 mm (-) - mm Figure  A.2.3    Cyclic test result for test combination CLT-B-N-L-C04 127    MECHANICAL PROPERTIES Fmax (+) 25.14 kN (-) - kN dFmax  (+) 31.30 mm (-) - mm 0.1 Fmax (+) 2.51 kN (-) - kN 0.4 Fmax (+) 10.05 kN (-) - kN 0.8 Fmax (+) 20.11 kN (-) - kN D (+) 5.73 [-] (-) - [-] Ke (+) 2.93 kN/mm (-) - kN/mm Fy  CUREE (+) 21.83 kN (-) - kN ∆y  CUREE (+) 7.45 mm (-) - mm Figure  A.2.4    Cyclic test result for test combination CLT-B-N-L-C05 128   MECHANICAL PROPERTIES Fmax (+) 27.73 kN (-) -29.36 kN dFmax  (+) 32.86 mm (-) -16.55 mm 0.1 Fmax (+) 2.77 kN (-) -2.94 kN 0.4 Fmax (+) 11.09 kN (-) -11.74 kN 0.8 Fmax (+) 22.18 kN (-) -23.49 kN D (+) 3.88 [-] (-) 6.92 [-] Ke (+) 2.17 kN/mm (-) 4.57 kN/mm Fy  CUREE (+) 23.93 kN (-) -25.69 kN ∆y  CUREE (+) 11.02 mm (-) -5.62 mm Figure  A.2.5    Cyclic test result for test combination CLT-B-N-P-C02 129    MECHANICAL PROPERTIES Fmax (+) 24.96 kN (-) -29.36 kN dFmax  (+) 31.83 mm (-) -18.02 mm 0.1 Fmax (+) 2.50 kN (-) -2.94 kN 0.4 Fmax (+) 9.98 kN (-) -11.74 kN 0.8 Fmax (+) 19.97 kN (-) -23.48 kN D (+) 3.95 [-] (-) 6.30 [-] Ke (+) 2.13 kN/mm (-) 4.11 kN/mm Fy  CUREE (+) 21.79 kN (-) -24.96 kN ∆y  CUREE (+) 10.23 mm (-) -6.08 mm Figure  A.2.6    Cyclic test result for test combination CLT-B-N-P-C03 130    MECHANICAL PROPERTIES Fmax (+) 23.72 kN (-) -26.00 kN dFmax  (+) 28.51 mm (-) -17.82 mm 0.1 Fmax (+) 2.37 kN (-) -2.60 kN 0.4 Fmax (+) 9.49 kN (-) -10.40 kN 0.8 Fmax (+) 18.98 kN (-) -20.80 kN D (+) 4.85 [-] (-) 4.09 [-] Ke (+) 2.48 kN/mm (-) 2.39 kN/mm Fy  CUREE (+) 20.59 kN (-) -22.87 kN ∆y  CUREE (+) 8.31 mm (-) -9.59 mm Figure  A.2.7    Cyclic test result for test combination CLT-B-N-P-C04 131    MECHANICAL PROPERTIES Fmax (+) 25.94 kN (-) -35.83 kN dFmax  (+) 29.15 mm (-) -32.81 mm 0.1 Fmax (+) 2.59 kN (-) -3.58 kN 0.4 Fmax (+) 10.38 kN (-) -14.33 kN 0.8 Fmax (+) 20.75 kN (-) -28.67 kN D (+) 3.55 [-] (-) 6.31 [-] Ke (+) 2.19 kN/mm (-) 4.24 kN/mm Fy  CUREE (+) 22.26 kN (-) -31.80 kN ∆y  CUREE (+) 10.16 mm (-) -7.49 mm Figure  A.2.8    Cyclic test result for test combination CLT-B-N-P-C05 132    MECHANICAL PROPERTIES Fmax (+) 21.96 kN (-) -34.45 kN dFmax  (+) 29.20 mm (-) -31.98 mm 0.1 Fmax (+) 2.20 kN (-) -3.45 kN 0.4 Fmax (+) 8.78 kN (-) -13.78 kN 0.8 Fmax (+) 17.57 kN (-) -27.56 kN D (+) 5.84 [-] (-) 7.26 [-] Ke (+) 2.82 kN/mm (-) 5.36 kN/mm Fy  CUREE (+) 19.05 kN (-) -30.40 kN ∆y  CUREE (+) 6.75 mm (-) -5.67 mm Figure  A.2.9    Cyclic test result for test combination CLT-B-N-P-C07 133 A.3 Pictures  Figure  A.3.1    Bracket B under monotonic loading  Figure  A.3.2    Pull-out failure of Bracket B under cyclic loading  134 Figure  A.3.3    Bracket B in test set-up for testing perpendicular to the grain   Figure  A.3.4    Fracture failure of Bracket B under cyclic loading: tension loading (left), compression loading (right)     135Appendix B  Bracket A with 12 Ring Shank Nails 3.76 mm x 76 mm B.1 Monotonic Test Results   Test ID   A-R-L-M5 A-R-L-M7 A-R-L-M10 Average Fmax [kN] 38.3 41.4 43.0 38.8 dFmax [mm] 14.0 15.0 23.0 15.0 Fult [kN] 30.6 33.1 34.4 31.0 ∆ult [mm] 22.0 23.0 29.0 25.3 D  [-]       4.1 Ke [kN/mm]       5.7 Fy - CUREE [kN]       34.7 ∆y - CUREE [mm]       6.1  Figure  B.1.1    Monotonic test results for connection combination A-R-L-M 136    Test ID   A-R-P-M5 A-R-P-M6 A-R-P-M9 Average Fmax [kN] 46.0 45.1 42.9 44.6 dFmax [mm] 25.0 23.0 24.0 24.0 Fult [kN] 36.8 36.1 34.3 35.7 ∆ult [mm] 38.0 34.0 29.0 33.3 D  [-]       3.3 Ke [kN/mm]       4.0 Fy - CUREE [kN]       39.7 ∆y - CUREE [mm]       10.0  Figure  B.1.2    Monotonic test results for connection combination A-R-P-M  137B.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 42.23 kN (-) - kN dFmax  (+) 14.42 mm (-) - mm 0.1 Fmax (+) 4.22 kN (-) - kN 0.4 Fmax (+) 16.89 kN (-) - kN 0.8 Fmax (+) 33.79 kN (-) - kN D (+) 5.43 [-] (-) - [-] Ke (+) 8.91 kN/mm (-) - kN/mm Fy  CUREE (+) 34.26 kN (-) - kN ∆y  CUREE (+) 3.85 mm (-) - mm Figure  B.2.1    Cyclic test result for test combination CLT-A-R-L-C04 138    MECHANICAL PROPERTIES Fmax (+) 45.06 kN (-) - kN dFmax  (+) 14.55 mm (-) - mm 0.1 Fmax (+) 4.51 kN (-) - kN 0.4 Fmax (+) 18.02 kN (-) - kN 0.8 Fmax (+) 36.05 kN (-) - kN D (+) 5.25 [-] (-) - [-] Ke (+) 10.27 kN/mm (-) - kN/mm Fy  CUREE (+) 37.85 kN (-) - kN ∆y  CUREE (+) 3.69 mm (-) - mm Figure  B.2.2    Cyclic test result for test combination CLT-A-R-L-C05 139    MECHANICAL PROPERTIES Fmax (+) 45.23 kN (-) - kN dFmax  (+) 14.50 mm (-) - mm 0.1 Fmax (+) 4.52 kN (-) - kN 0.4 Fmax (+) 18.09 kN (-) - kN 0.8 Fmax (+) 36.18 kN (-) - kN D (+) 6.19 [-] (-) - [-] Ke (+) 9.25 kN/mm (-) - kN/mm Fy  CUREE (+) 39.66 kN (-) - kN ∆y  CUREE (+) 4.29 mm (-) - mm Figure  B.2.3    Cyclic test result for test combination CLT-A-R-L-C06 140    MECHANICAL PROPERTIES Fmax (+) 37.49 kN (-) - kN dFmax  (+) 20.75 mm (-) - mm 0.1 Fmax (+) 3.75 kN (-) - kN 0.4 Fmax (+) 15.00 kN (-) - kN 0.8 Fmax (+) 29.99 kN (-) - kN D (+) 5.83 [-] (-) - [-] Ke (+) 5.41 kN/mm (-) - kN/mm Fy  CUREE (+) 32.16 kN (-) - kN ∆y  CUREE (+) 5.94 mm (-) - mm Figure  B.2.4    Cyclic test result for test combination CLT-A-R-L-C07 141    MECHANICAL PROPERTIES Fmax (+) 40.52 kN (-) - kN dFmax  (+) 22.61 mm (-) - mm 0.1 Fmax (+) 4.05 kN (-) - kN 0.4 Fmax (+) 16.21 kN (-) - kN 0.8 Fmax (+) 32.42 kN (-) - kN D (+) 4.52 [-] (-) - [-] Ke (+) 4.86 kN/mm (-) - kN/mm Fy  CUREE (+) 35.04 kN (-) - kN ∆y  CUREE (+) 7.20 mm (-) - mm Figure  B.2.5    Cyclic test result for test combination CLT-A-R-L-C08 142    MECHANICAL PROPERTIES Fmax (+) 43.38 kN (-) -40.91 kN dFmax  (+) 24.27 mm (-) -17.58 mm 0.1 Fmax (+) 4.34 kN (-) -4.09 kN 0.4 Fmax (+) 17.35 kN (-) -16.37 kN 0.8 Fmax (+) 34.70 kN (-) -32.73 kN D (+) 4.08 [-] (-) 5.93 [-] Ke (+) 4.90 kN/mm (-) 7.67 kN/mm Fy  CUREE (+) 36.11 kN (-) -35.70 kN ∆y  CUREE (+) 7.36 mm (-) -4.65 mm Figure  B.2.6    Cyclic test result for test combination CLT-A-R-P-C01 143    MECHANICAL PROPERTIES Fmax (+) 41.35 kN (-) -45.98 kN dFmax  (+) 23.14 mm (-) -14.94 mm 0.1 Fmax (+) 4.14 kN (-) -4.60 kN 0.4 Fmax (+) 16.54 kN (-) -18.39 kN 0.8 Fmax (+) 33.08 kN (-) -36.78 kN D (+) 4.51 [-] (-) 3.27 [-] Ke (+) 5.67 kN/mm (-) 5.21 kN/mm Fy  CUREE (+) 34.20 kN (-) -39.30 kN ∆y  CUREE (+) 6.03 mm (-) -7.54 mm Figure  B.2.7    Cyclic test result for test combination CLT-A-R-P-C02 144    MECHANICAL PROPERTIES Fmax (+) 42.16 kN (-) -49.66 kN dFmax  (+) 24.46 mm (-) -14.60 mm 0.1 Fmax (+) 4.22 kN (-) -4.97 kN 0.4 Fmax (+) 16.86 kN (-) -19.86 kN 0.8 Fmax (+) 33.73 kN (-) -39.73 kN D (+) 3.78 [-] (-) 4.09 [-] Ke (+) 4.91 kN/mm (-) 6.85 kN/mm Fy  CUREE (+) 36.11 kN (-) -39.80 kN ∆y  CUREE (+) 7.35 mm (-) -5.81 mm Figure  B.2.8    Cyclic test result for test combination CLT-A-R-P-C06 145    MECHANICAL PROPERTIES Fmax (+) 47.43 kN (-) -56.64 kN dFmax  (+) 28.42 mm (-) -26.76 mm 0.1 Fmax (+) 4.74 kN (-) -5.66 kN 0.4 Fmax (+) 18.97 kN (-) -22.66 kN 0.8 Fmax (+) 37.94 kN (-) -45.31 kN D (+) 3.62 [-] (-) 2.92 [-] Ke (+) 4.41 kN/mm (-) 4.80 kN/mm Fy  CUREE (+) 41.55 kN (-) -49.01 kN ∆y  CUREE (+) 9.42 mm (-) -10.21 mm Figure  B.2.9    Cyclic test result for test combination CLT-A-R-P-C07 146    MECHANICAL PROPERTIES Fmax (+) 41.98 kN (-) -51.04 kN dFmax  (+) 28.17 mm (-) -27.24 mm 0.1 Fmax (+) 4.20 kN (-) -5.10 kN 0.4 Fmax (+) 16.79 kN (-) -20.42 kN 0.8 Fmax (+) 33.58 kN (-) -40.83 kN D (+) 4.81 [-] (-) 3.01 [-] Ke (+) 4.77 kN/mm (-) 3.88 kN/mm Fy  CUREE (+) 36.45 kN (-) -45.04 kN ∆y  CUREE (+) 7.65 mm (-) -11.61 mm Figure  B.2.10    Cyclic test result for test combination CLT-A-R-P-C08 147 B.3 Pictures  Figure  B.3.1    Bracket A with 12 Ring Shank Nails under monotonic loading   Figure  B.3.2    Bracket A under cyclic loading: one nail failed in shear, one nail failed in pull-out   148     Figure  B.3.3    Four stages of a cyclic loading with Bracket A  149  Figure  B.3.4    Bracket A under cyclic loading perpendicular to the grain  150Appendix C  Bracket A with 12 Ring Shank Nails 4.2 mm x 60 mm C.1 Monotonic Test Results   Test ID   A-r-L-M4 A-r-L-M6 A-r-L-M9 Average Fmax [kN] 40.6 34.4 40.6 38.1 dFmax [mm] 15.0 12.0 17.0 16.0 Fult [kN] 32.5 27.5 32.5 30.4 ∆ult [mm] 22.0 25.0 27.0 23.0 D  [-]       4.0 Ke [kN/mm]       6.1 Fy - CUREE [kN]       34.7 ∆y - CUREE [mm]       5.7  Figure  C.1.1    Monotonic test results for connection combination A-r-L-M  151   Test ID   A-r-P-M7 A-r-P-M8 A-r-P-M10 Average Fmax [kN] 35.1 40.2 40.5 38.6 dFmax [mm] 19.0 18.0 18.0 18.0 Fult [kN] 28.1 32.2 32.4 30.9 ∆ult [mm] 29.0 27.0 27.0 27.1 D  [-]       4.3 Ke [kN/mm]       5.5 Fy - CUREE [kN]       34.8 ∆y - CUREE [mm]       6.3 Figure  C.1.2    Monotonic test results for connection combination A-r-P-M  152C.2 Cyclic Test    MECHANICAL PROPERTIES Fmax (+) 35.41 kN (-) - kN dFmax  (+) 10.60 mm (-) - mm 0.1 Fmax (+) 3.54 kN (-) - kN 0.4 Fmax (+) 14.16 kN (-) - kN 0.8 Fmax (+) 28.33 kN (-) - kN D (+) 5.24 [-] (-) - [-] Ke (+) 8.36 kN/mm (-) - kN/mm Fy  CUREE (+) 31.37 kN (-) - kN ∆y  CUREE (+) 3.75 mm (-) - mm Figure  C.2.1    Cyclic test result for test combination CLT-A-r-L-C01 153  MECHANICAL PROPERTIES Fmax (+) 35.44 kN (-) - kN dFmax  (+) 11.46 mm (-) - mm 0.1 Fmax (+) 3.54 kN (-) - kN 0.4 Fmax (+) 14.18 kN (-) - kN 0.8 Fmax (+) 28.35 kN (-) - kN D (+) 8.86 [-] (-) - [-] Ke (+) 10.21 kN/mm (-) - kN/mm Fy  CUREE (+) 31.48 kN (-) - kN ∆y  CUREE (+) 3.08 mm (-) - mm Figure  C.2.2    Cyclic test result for test combination CLT-A-r-L-C02 154    MECHANICAL PROPERTIES Fmax (+) 32.92 kN (-) - kN dFmax  (+) 11.76 mm (-) - mm 0.1 Fmax (+) 3.29 kN (-) - kN 0.4 Fmax (+) 13.17 kN (-) - kN 0.8 Fmax (+) 26.34 kN (-) - kN D (+) 5.44 [-] (-) - [-] Ke (+) 6.80 kN/mm (-) - kN/mm Fy  CUREE (+) 29.76 kN (-) - kN ∆y  CUREE (+) 4.38 mm (-) - mm Figure  C.2.3    Cyclic test result for test combination CLT-A-r-L-C03 155    MECHANICAL PROPERTIES Fmax (+) 35.86 kN (-) - kN dFmax  (+) 15.72 mm (-) - mm 0.1 Fmax (+) 3.59 kN (-) - kN 0.4 Fmax (+) 14.35 kN (-) - kN 0.8 Fmax (+) 28.69 kN (-) - kN D (+) 4.45 [-] (-) - [-] Ke (+) 5.89 kN/mm (-) - kN/mm Fy  CUREE (+) 31.26 kN (-) - kN ∆y  CUREE (+) 5.31 mm (-) - mm Figure  C.2.4    Cyclic test result for test combination CLT-A-r-L-C04 156    MECHANICAL PROPERTIES Fmax (+) 34.42 kN (-) - kN dFmax  (+) 18.46 mm (-) - mm 0.1 Fmax (+) 3.44 kN (-) - kN 0.4 Fmax (+) 13.77 kN (-) - kN 0.8 Fmax (+) 27.54 kN (-) - kN D (+) 4.19 [-] (-) - [-] Ke (+) 4.65 kN/mm (-) - kN/mm Fy  CUREE (+) 30.61 kN (-) - kN ∆y  CUREE (+) 6.58 mm (-) - mm Figure  C.2.5    Cyclic test result for test combination CLT-A-r-L-C05 157    MECHANICAL PROPERTIES Fmax (+) 39.83 kN (-) -48.95 kN dFmax  (+) 18.60 mm (-) -20.75 mm 0.1 Fmax (+) 3.98 kN (-) -4.89 kN 0.4 Fmax (+) 15.93 kN (-) -19.58 kN 0.8 Fmax (+) 31.86 kN (-) -39.16 kN D (+) 6.54 [-] (-) 4.70 [-] Ke (+) 6.62 kN/mm (-) 5.94 kN/mm Fy  CUREE (+) 34.94 kN (-) -43.09 kN ∆y  CUREE (+) 5.27 mm (-) -7.26 mm Figure  C.2.6    Cyclic test result for test combination CLT-A-r-P-C03 158    MECHANICAL PROPERTIES Fmax (+) 37.24 kN (-) -40.39 kN dFmax  (+) 20.12 mm (-) -21.38 mm 0.1 Fmax (+) 3.72 kN (-) -4.04 kN 0.4 Fmax (+) 14.90 kN (-) -16.16 kN 0.8 Fmax (+) 29.79 kN (-) -32.31 kN D (+) 5.63 [-] (-) 3.10 [-] Ke (+) 6.20 kN/mm (-) 4.24 kN/mm Fy  CUREE (+) 31.71 kN (-) -34.26 kN ∆y  CUREE (+) 5.11 mm (-) -8.08 mm Figure  C.2.7    Cyclic test result for test combination CLT-A-r-P-C04 159    MECHANICAL PROPERTIES Fmax (+) 41.22 kN (-) -41.30 kN dFmax  (+) 21.53 mm (-) -21.87 mm 0.1 Fmax (+) 4.12 kN (-) -4.13 kN 0.4 Fmax (+) 16.49 kN (-) -16.52 kN 0.8 Fmax (+) 32.98 kN (-) -33.04 kN D (+) 4.37 [-] (-) 5.53 [-] Ke (+) 5.91 kN/mm (-) 7.30 kN/mm Fy  CUREE (+) 34.82 kN (-) -35.89 kN ∆y  CUREE (+) 5.89 mm (-) -4.92 mm Figure  C.2.8    Cyclic test result for test combination CLT-A-r-P-C05 160    MECHANICAL PROPERTIES Fmax (+) 41.39 kN (-) -47.80 kN dFmax  (+) 20.65 mm (-) -21.53 mm 0.1 Fmax (+) 4.14 kN (-) -4.78 kN 0.4 Fmax (+) 16.56 kN (-) -19.12 kN 0.8 Fmax (+) 33.11 kN (-) -38.24 kN D (+) 4.27 [-] (-) 3.26 [-] Ke (+) 5.23 kN/mm (-) 4.97 kN/mm Fy  CUREE (+) 36.62 kN (-) -41.58 kN ∆y  CUREE (+) 7.00 mm (-) -8.37 mm Figure  C.2.9    Cyclic test result for test combination CLT-A-r-P-C06 161    MECHANICAL PROPERTIES Fmax (+) 41.22 kN (-) -47.90 kN dFmax  (+) 21.68 mm (-) -21.53 mm 0.1 Fmax (+) 4.12 kN (-) -4.79 kN 0.4 Fmax (+) 16.49 kN (-) -19.16 kN 0.8 Fmax (+) 32.98 kN (-) -38.32 kN D (+) 3.21 [-] (-) 3.01 [-] Ke (+) 4.04 kN/mm (-) 4.11 kN/mm Fy  CUREE (+) 36.29 kN (-) -42.59 kN ∆y  CUREE (+) 8.99 mm (-) -10.36 mm Figure  C.2.10    Cyclic test result for test combination CLT-A-r-P-C07 162 C.3 Pictures  Figure  C.3.1    Front view of Bracket A with short Rink Shank Nails  Figure  C.3.2    Pull-out failure of Bracket A with small Ring Shank Nails under cyclic loading  163 Figure  C.3.3    Front view of test set-up perpendicular to the grain  Figure  C.3.4    Destructed CLT panel after test perpendicular to the grain  164Appendix D  Bracket A with 9 Screws 5 mm x 90 mm D.1 Monotonic Test Results   Test ID   A-S-L-M1 A-S-L-M2 A-S-L-M3 Average Fmax [kN] 40.1 38.0 38.9 38.7 dFmax [mm] 17.0 16.0 19.0 18.0 Fult [kN] 32.1 30.4 31.1 31.0 ∆ult [mm] 26.0 25.0 25.0 25.0 D  [-]       3.6 Ke [kN/mm]       4.9 Fy - CUREE [kN]       34.1 ∆y - CUREE [mm]       6.9 Figure  D.1.1    Monotonic test results for connection combination A-S-L-M 165    Test ID   A-S-P-M7 A-S-P-M8 A-S-P-M0 Average Fmax [kN] 49.4 44.7 44.4 45.6 dFmax [mm] 23.0 25.0 29.0 23.0 Fult [kN] 39.5 35.7 35.5 36.5 ∆ult [mm] 32.0 34.0 37.0 33.8 D  [-]       3.9 Ke [kN/mm]       4.7 Fy - CUREE [kN]       41.0 ∆y - CUREE [mm]       8.7 Figure  D.1.2    Monotonic test results for connection combination A-S-P-M  166D.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 52.00 kN (-) - kN dFmax  (+) 17.33 mm (-) - mm 0.1 Fmax (+) 5.20 kN (-) - kN 0.4 Fmax (+) 20.80 kN (-) - kN 0.8 Fmax (+) 41.60 kN (-) - kN D (+) 3.02 [-] (-) - [-] Ke (+) 5.90 kN/mm (-) - kN/mm Fy  CUREE (+) 44.73 kN (-) - kN ∆y  CUREE (+) 7.58 mm (-) - mm Figure  D.2.1    Cyclic test result for test combination CLT-A-S-L-C07 167    MECHANICAL PROPERTIES Fmax (+) 42.44 kN (-) - kN dFmax  (+) 17.75 mm (-) - mm 0.1 Fmax (+) 4.24 kN (-) - kN 0.4 Fmax (+) 16.97 kN (-) - kN 0.8 Fmax (+) 33.95 kN (-) - kN D (+) 3.35 [-] (-) - [-] Ke (+) 4.83 kN/mm (-) - kN/mm Fy  CUREE (+) 37.15 kN (-) - kN ∆y  CUREE (+) 7.68 mm (-) - mm Figure  D.2.2    Cyclic test result for test combination CLT-A-S-L-C08 168    MECHANICAL PROPERTIES Fmax (+) 43.63 kN (-) - kN dFmax  (+) 15.89 mm (-) - mm 0.1 Fmax (+) 4.36 kN (-) - kN 0.4 Fmax (+) 17.45 kN (-) - kN 0.8 Fmax (+) 34.90 kN (-) - kN D (+) 3.19 [-] (-) - [-] Ke (+) 4.90 kN/mm (-) - kN/mm Fy  CUREE (+) 37.99 kN (-) - kN ∆y  CUREE (+) 7.75 mm (-) - mm Figure  D.2.3    Cyclic test result for test combination CLT-A-S-L-C09 169    MECHANICAL PROPERTIES Fmax (+) 51.02 kN (-) - kN dFmax  (+) 26.71 mm (-) - mm 0.1 Fmax (+) 5.10 kN (-) - kN 0.4 Fmax (+) 20.41 kN (-) - kN 0.8 Fmax (+) 40.82 kN (-) - kN D (+) 3.61 [-] (-) - [-] Ke (+) 4.66 kN/mm (-) - kN/mm Fy  CUREE (+) 43.85 kN (-) - kN ∆y  CUREE (+) 9.41 mm (-) - mm Figure  D.2.4    Cyclic test result for test combination CLT-A-S-L-C10 170    MECHANICAL PROPERTIES Fmax (+) 40.70 kN (-) - kN dFmax  (+) 21.68 mm (-) - mm 0.1 Fmax (+) 4.07 kN (-) - kN 0.4 Fmax (+) 16.28 kN (-) - kN 0.8 Fmax (+) 32.56 kN (-) - kN D (+) 4.16 [-] (-) - [-] Ke (+) 4.99 kN/mm (-) - kN/mm Fy  CUREE (+) 34.77 kN (-) - kN ∆y  CUREE (+) 6.97 mm (-) - mm Figure  D.2.5    Cyclic test result for test combination CLT-A-S-L-C 171    MECHANICAL PROPERTIES Fmax (+) 48.08 kN (-) -53.02 kN dFmax  (+) 24.46 mm (-) -19.53 mm 0.1 Fmax (+) 4.81 kN (-) -5.30 kN 0.4 Fmax (+) 19.23 kN (-) -21.21 kN 0.8 Fmax (+) 38.47 kN (-) -42.42 kN D (+) 3.12 [-] (-) 3.65 [-] Ke (+) 3.49 kN/mm (-) 5.75 kN/mm Fy  CUREE (+) 42.64 kN (-) -45.55 kN ∆y  CUREE (+) 12.20 mm (-) -7.93 mm Figure  D.2.6    Cyclic test result for test combination CLT-A-S-P-C05 172    MECHANICAL PROPERTIES Fmax (+) 44.92 kN (-) -54.25 kN dFmax  (+) 22.85 mm (-) -23.73 mm 0.1 Fmax (+) 4.49 kN (-) -5.42 kN 0.4 Fmax (+) 17.97 kN (-) -21.70 kN 0.8 Fmax (+) 35.94 kN (-) -43.40 kN D (+) 4.08 [-] (-) 3.70 [-] Ke (+) 4.45 kN/mm (-) 5.75 kN/mm Fy  CUREE (+) 38.12 kN (-) -45.70 kN ∆y  CUREE (+) 8.56 mm (-) -7.95 mm Figure  D.2.7    Cyclic test result for test combination CLT-A-S-P-C06 173    MECHANICAL PROPERTIES Fmax (+) 48.26 kN (-) -52.63 kN dFmax  (+) 22.21 mm (-) -24.46 mm 0.1 Fmax (+) 4.83 kN (-) -5.26 kN 0.4 Fmax (+) 19.30 kN (-) -21.05 kN 0.8 Fmax (+) 38.61 kN (-) -42.10 kN D (+) 4.27 [-] (-) 3.88 [-] Ke (+) 4.90 kN/mm (-) 6.15 kN/mm Fy  CUREE (+) 40.75 kN (-) -44.86 kN ∆y  CUREE (+) 8.32 mm (-) -7.30 mm Figure  D.2.8    Cyclic test result for test combination CLT-A-S-P-C08 174    MECHANICAL PROPERTIES Fmax (+) 45.28 kN (-) -55.57 kN dFmax  (+) 28.85 mm (-) -26.51 mm 0.1 Fmax (+) 4.53 kN (-) -5.56 kN 0.4 Fmax (+) 18.11 kN (-) -22.23 kN 0.8 Fmax (+) 36.22 kN (-) -44.46 kN D (+) 4.07 [-] (-) 3.04 [-] Ke (+) 4.44 kN/mm (-) 4.68 kN/mm Fy  CUREE (+) 37.59 kN (-) -48.22 kN ∆y  CUREE (+) 8.46 mm (-) -10.31 mm Figure  D.2.9    Cyclic test result for test combination CLT-A-S-P-C09 175    MECHANICAL PROPERTIES Fmax (+) 24.50 kN (-) - kN dFmax  (+) 17.66 mm (-) - mm 0.1 Fmax (+) 2.45 kN (-) - kN 0.4 Fmax (+) 9.80 kN (-) - kN 0.8 Fmax (+) 19.60 kN (-) - kN D (+) 5.84 [-] (-) - [-] Ke (+) 4.82 kN/mm (-) - kN/mm Fy  CUREE (+) 22.70 kN (-) - kN ∆y  CUREE (+) 4.71 mm (-) - mm Figure  D.2.10    Cyclic test result for test combination CLT-A-S-P-C10 176 D.3 Pictures  Figure  D.3.1    Bracket A with 9 Screws 5 x 90 mm under monotonic loading  Figure  D.3.2    Deformed connector Bracket A under cyclic loading  177 Figure  D.3.3    Cracked CLT panel under cyclic loading perpendicular to the grain  Figure  D.3.4    Shear failure of all nine screws under cyclic loading  178Appendix E  Bracket A with 18 Screws 4 mm x 70 mm E.1 Monotonic Test Results   Test ID   A-s-L-M4 A-s-L-M5 A-s-L-M6 Average Fmax [kN] 52.1 44.7 49.2 49.7 dFmax [mm] 23.0 25.0 21.0 21.0 Fult [kN] 41.7 35.7 39.3 39.8 ∆ult [mm] 31.0 28.0 28.0 27.9 D  [-]       3.1 Ke [kN/mm]       4.9 Fy - CUREE [kN]       44.0 ∆y - CUREE [mm]       9.0 Figure  E.1.1    Monotonic test results for connection combination A-s-L-M  179   Test ID   A-s-P-M5 A-s-P-M6 A-s-P-M9 Average Fmax [kN] 45.7 48.4 47.5 46.1 dFmax [mm] 19.0 25.0 34.0 25.0 Fult [kN] 36.5 38.7 38.0 36.9 ∆ult [mm] 32.0 35.0 40.0 34.7 D  [-]       3.6 Ke [kN/mm]       4.3 Fy - CUREE [kN]       42.2 ∆y - CUREE [mm]       9.7 Figure  E.1.2    Monotonic test results for connection combination A-S-L-M  180E.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 53.93 kN (-) - kN dFmax  (+) 16.01 mm (-) - mm 0.1 Fmax (+) 5.39 kN (-) - kN 0.4 Fmax (+) 21.57 kN (-) - kN 0.8 Fmax (+) 43.14 kN (-) - kN D (+) 4.05 [-] (-) - [-] Ke (+) 6.82 kN/mm (-) - kN/mm Fy  CUREE (+) 47.01 kN (-) - kN ∆y  CUREE (+) 6.89 mm (-) - mm Figure  E.2.1    Cyclic test result for test combination CLT-A-s-L-C04 181    MECHANICAL PROPERTIES Fmax (+) 45.71 kN (-) - kN dFmax  (+) 15.53 mm (-) - mm 0.1 Fmax (+) 4.57 kN (-) - kN 0.4 Fmax (+) 18.28 kN (-) - kN 0.8 Fmax (+) 36.57 kN (-) - kN D (+) 3.51 [-] (-) - [-] Ke (+) 6.37 kN/mm (-) - kN/mm Fy  CUREE (+) 38.32 kN (-) - kN ∆y  CUREE (+) 6.01 mm (-) - mm Figure  E.2.2    Cyclic test result for test combination CLT-A-s-L-C05 182    MECHANICAL PROPERTIES Fmax (+) 55.00 kN (-) - kN dFmax  (+) 21.32 mm (-) - mm 0.1 Fmax (+) 5.50 kN (-) - kN 0.4 Fmax (+) 22.00 kN (-) - kN 0.8 Fmax (+) 44.00 kN (-) - kN D (+) 3.44 [-] (-) - [-] Ke (+) 5.97 kN/mm (-) - kN/mm Fy  CUREE (+) 46.71 kN (-) - kN ∆y  CUREE (+) 7.82 mm (-) - mm Figure  E.2.3    Cyclic test result for test combination CLT-A-s-L-C06 183    MECHANICAL PROPERTIES Fmax (+) 51.90 kN (-) - kN dFmax  (+) 20.36 mm (-) - mm 0.1 Fmax (+) 5.19 kN (-) - kN 0.4 Fmax (+) 20.76 kN (-) - kN 0.8 Fmax (+) 41.52 kN (-) - kN D (+) 3.54 [-] (-) - [-] Ke (+) 5.66 kN/mm (-) - kN/mm Fy  CUREE (+) 44.06 kN (-) - kN ∆y  CUREE (+) 7.78 mm (-) - mm Figure  E.2.4    Cyclic test result for test combination CLT-A-s-L-C07 184    MECHANICAL PROPERTIES Fmax (+) 34.24 kN (-) - kN dFmax  (+) 19.53 mm (-) - mm 0.1 Fmax (+) 3.42 kN (-) - kN 0.4 Fmax (+) 13.69 kN (-) - kN 0.8 Fmax (+) 27.39 kN (-) - kN D (+) 2.32 [-] (-) - [-] Ke (+) 2.72 kN/mm (-) - kN/mm Fy  CUREE (+) 27.14 kN (-) - kN ∆y  CUREE (+) 9.98 mm (-) - mm Figure  E.2.5    Cyclic test result for test combination CLT-A-s-L-C08 185    MECHANICAL PROPERTIES Fmax (+) 51.16 kN (-) -58.34 kN dFmax  (+) 26.07 mm (-) -20.16 mm 0.1 Fmax (+) 5.12 kN (-) -5.83 kN 0.4 Fmax (+) 20.46 kN (-) -23.34 kN 0.8 Fmax (+) 40.93 kN (-) -46.67 kN D (+) 3.21 [-] (-) 4.33 [-] Ke (+) 3.54 kN/mm (-) 8.07 kN/mm Fy  CUREE (+) 44.69 kN (-) -48.67 kN ∆y  CUREE (+) 12.63 mm (-) -6.03 mm Figure  E.2.6    Cyclic test result for test combination CLT-A-s-P-C03 186    MECHANICAL PROPERTIES Fmax (+) 49.69 kN (-) -60.53 kN dFmax  (+) 24.66 mm (-) -27.10 mm 0.1 Fmax (+) 4.97 kN (-) -6.05 kN 0.4 Fmax (+) 19.88 kN (-) -24.21 kN 0.8 Fmax (+) 39.76 kN (-) -48.43 kN D (+) 4.23 [-] (-) 3.23 [-] Ke (+) 4.79 kN/mm (-) 5.47 kN/mm Fy  CUREE (+) 42.81 kN (-) -50.52 kN ∆y  CUREE (+) 8.95 mm (-) -9.24 mm Figure  E.2.7    Cyclic test result for test combination CLT-A-s-P-C04 187    MECHANICAL PROPERTIES Fmax (+) 52.00 kN (-) -64.80 kN dFmax  (+) 36.76 mm (-) -26.80 mm 0.1 Fmax (+) 5.20 kN (-) -6.48 kN 0.4 Fmax (+) 20.80 kN (-) -25.92 kN 0.8 Fmax (+) 41.60 kN (-) -51.84 kN D (+) 3.39 [-] (-) 3.11 [-] Ke (+) 4.00 kN/mm (-) 5.05 kN/mm Fy  CUREE (+) 46.86 kN (-) -54.37 kN ∆y  CUREE (+) 11.71 mm (-) -10.78 mm Figure  E.2.8    Cyclic test result for test combination CLT-A-s-P-C07 188    MECHANICAL PROPERTIES Fmax (+) 51.30 kN (-) -63.15 kN dFmax  (+) 28.46 mm (-) -17.97 mm 0.1 Fmax (+) 5.13 kN (-) -6.31 kN 0.4 Fmax (+) 20.52 kN (-) -25.26 kN 0.8 Fmax (+) 41.04 kN (-) -50.52 kN D (+) 4.57 [-] (-) 3.79 [-] Ke (+) 6.46 kN/mm (-) 6.79 kN/mm Fy  CUREE (+) 44.20 kN (-) -56.62 kN ∆y  CUREE (+) 6.84 mm (-) -8.34 mm Figure  E.2.9    Cyclic test result for test combination CLT-A-s-P-C08 189    MECHANICAL PROPERTIES Fmax (+) 47.61 kN (-) -62.44 kN dFmax  (+) 23.14 mm (-) -20.31 mm 0.1 Fmax (+) 4.76 kN (-) -6.24 kN 0.4 Fmax (+) 19.04 kN (-) -24.97 kN 0.8 Fmax (+) 38.09 kN (-) -49.95 kN D (+) 4.08 [-] (-) 3.82 [-] Ke (+) 5.69 kN/mm (-) 6.28 kN/mm Fy  CUREE (+) 42.50 kN (-) -52.78 kN ∆y  CUREE (+) 7.46 mm (-) -8.41 mm Figure  E.2.10    Cyclic test result for test combination CLT-A-s-P-C09 190 E.3 Pictures  Figure  E.3.1    Bracket A with 18 screws 4 x 70 mm under monotonic loading (front view)  Figure  E.3.2    Bracket A with 18 screws 4 x 70 mm under monotonic loading (side view)  191 Figure  E.3.3    series of four stages of Bracket A under cyclic loading parallel to the grain  192 Figure  E.3.4    Front view of Bracket A with screws 4 x 70 mm under cyclic loading perpendicular to the grain  Figure  E.3.5    destructed CLT panel under cyclic loading perpendicular to the grain  193Appendix F  Bracket A with 18 Spiral Nails 4.2 mm x 89 mm F.1 Monotonic Test Results   Test ID   A-N-L-M5 A-N-L-M6 A-N-L-M7 Average Fmax [kN] 46.5 51.7 59.9 52.0 dFmax [mm] 19.0 21.0 26.0 21.0 Fult [kN] 37.2 41.4 47.9 41.6 ∆ult [mm] 29.0 33.0 0.0 32.3 D  [-]       3.1 Ke [kN/mm]       4.7 Fy - CUREE [kN]       49.1 ∆y - CUREE [mm]       10.4 Figure  F.1.1    Monotonic test results for connection combination A-N-L-M  194    Test ID   A-N-P-M4 A-N-P-M5 A-N-P-M6 Average Fmax [kN] 50.4 51.5 54.5 51.9 dFmax [mm] 36.0 34.0 39.0 37.0 Fult [kN] 40.4 41.2 43.6 41.6 ∆ult [mm] 52.0 0.0 44.0 58.2 D  [-]       6.4 Ke [kN/mm]       5.1 Fy - CUREE [kN]       46.6 ∆y - CUREE [mm]       9.1 Figure  F.1.2    Monotonic test results for connection combination A-N-P-M  195F.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 51.00 kN (-) - kN dFmax  (+) 20.94 mm (-) - mm 0.1 Fmax (+) 5.10 kN (-) - kN 0.4 Fmax (+) 20.40 kN (-) - kN 0.8 Fmax (+) 40.80 kN (-) - kN D (+) 4.52 [-] (-) - [-] Ke (+) 6.26 kN/mm (-) - kN/mm Fy  CUREE (+) 46.37 kN (-) - kN ∆y  CUREE (+) 7.41 mm (-) - mm Figure  F.2.1    Cyclic test result for test combination CLT-A-N-L-C01 196    MECHANICAL PROPERTIES Fmax (+) 49.49 kN (-) - kN dFmax  (+) 19.53 mm (-) - mm 0.1 Fmax (+) 4.95 kN (-) - kN 0.4 Fmax (+) 19.80 kN (-) - kN 0.8 Fmax (+) 39.59 kN (-) - kN D (+) 7.16 [-] (-) - [-] Ke (+) 10.61 kN/mm (-) - kN/mm Fy  CUREE (+) 44.50 kN (-) - kN ∆y  CUREE (+) 4.19 mm (-) - mm Figure  F.2.2    Cyclic test result for test combination CLT-A-N-L-C02 197    MECHANICAL PROPERTIES Fmax (+) 44.87 kN (-) - kN dFmax  (+) 16.58 mm (-) - mm 0.1 Fmax (+) 4.49 kN (-) - kN 0.4 Fmax (+) 17.95 kN (-) - kN 0.8 Fmax (+) 35.89 kN (-) - kN D (+) 6.81 [-] (-) - [-] Ke (+) 10.12 kN/mm (-) - kN/mm Fy  CUREE (+) 40.91 kN (-) - kN ∆y  CUREE (+) 4.04 mm (-) - mm Figure  F.2.3    Cyclic test result for test combination CLT-A-N-L-C03 198    MECHANICAL PROPERTIES Fmax (+) 50.21 kN (-) - kN dFmax  (+) 16.14 mm (-) - mm 0.1 Fmax (+) 5.02 kN (-) - kN 0.4 Fmax (+) 20.08 kN (-) - kN 0.8 Fmax (+) 40.16 kN (-) - kN D (+) 5.68 [-] (-) - [-] Ke (+) 10.16 kN/mm (-) - kN/mm Fy  CUREE (+) 44.85 kN (-) - kN ∆y  CUREE (+) 4.41 mm (-) - mm Figure  F.2.4    Cyclic test result for test combination CLT-A-N-L-C04 199    MECHANICAL PROPERTIES Fmax (+) 48.03 kN (-) - kN dFmax  (+) 19.38 mm (-) - mm 0.1 Fmax (+) 4.80 kN (-) - kN 0.4 Fmax (+) 19.21 kN (-) - kN 0.8 Fmax (+) 38.43 kN (-) - kN D (+) 7.05 [-] (-) - [-] Ke (+) 9.69 kN/mm (-) - kN/mm Fy  CUREE (+) 43.33 kN (-) - kN ∆y  CUREE (+) 4.47 mm (-) - mm Figure  F.2.5    Cyclic test result for test combination CLT-A-N-L-C05 200    MECHANICAL PROPERTIES Fmax (+) 57.94 kN (-) - kN dFmax  (+) 35.30 mm (-) - mm 0.1 Fmax (+) 5.79 kN (-) - kN 0.4 Fmax (+) 23.18 kN (-) - kN 0.8 Fmax (+) 46.35 kN (-) - kN D (+) 5.30 [-] (-) - [-] Ke (+) 5.93 kN/mm (-) - kN/mm Fy  CUREE (+) 50.24 kN (-) - kN ∆y  CUREE (+) 8.48 mm (-) - mm Figure  F.2.6    Cyclic test result for test combination CLT-A-N-L-C06 201    MECHANICAL PROPERTIES Fmax (+) 44.86 kN (-) - kN dFmax  (+) 15.25 mm (-) - mm 0.1 Fmax (+) 4.49 kN (-) - kN 0.4 Fmax (+) 17.94 kN (-) - kN 0.8 Fmax (+) 35.89 kN (-) - kN D (+) 6.85 [-] (-) - [-] Ke (+) 8.17 kN/mm (-) - kN/mm Fy  CUREE (+) 40.57 kN (-) - kN ∆y  CUREE (+) 4.97 mm (-) - mm Figure  F.2.7    Cyclic test result for test combination CLT-A-N-L-C07 202    MECHANICAL PROPERTIES Fmax (+) 43.40 kN (-) -46.23 kN dFmax  (+) 17.87 mm (-) -19.72 mm 0.1 Fmax (+) 4.34 kN (-) -4.62 kN 0.4 Fmax (+) 17.36 kN (-) -18.49 kN 0.8 Fmax (+) 34.72 kN (-) -36.98 kN D (+) 5.19 [-] (-) 4.93 [-] Ke (+) 5.68 kN/mm (-) 6.75 kN/mm Fy  CUREE (+) 35.70 kN (-) -39.76 kN ∆y  CUREE (+) 6.28 mm (-) -5.89 mm Figure  F.2.8    Cyclic test result for test combination CLT-A-N-P-C02 203    MECHANICAL PROPERTIES Fmax (+) 48.58 kN (-) -49.65 kN dFmax  (+) 37.25 mm (-) -17.28 mm 0.1 Fmax (+) 4.86 kN (-) -4.96 kN 0.4 Fmax (+) 19.43 kN (-) -19.86 kN 0.8 Fmax (+) 38.86 kN (-) -39.72 kN D (+) 6.18 [-] (-) 4.15 [-] Ke (+) 6.23 kN/mm (-) 7.30 kN/mm Fy  CUREE (+) 44.13 kN (-) -42.75 kN ∆y  CUREE (+) 7.08 mm (-) -5.86 mm Figure  F.2.9    Cyclic test result for test combination CLT-A-N-P-C03 204    MECHANICAL PROPERTIES Fmax (+) 44.83 kN (-) -47.03 kN dFmax  (+) 23.92 mm (-) -19.77 mm 0.1 Fmax (+) 4.48 kN (-) -4.70 kN 0.4 Fmax (+) 17.93 kN (-) -18.81 kN 0.8 Fmax (+) 35.86 kN (-) -37.63 kN D (+) 4.57 [-] (-) 4.57 [-] Ke (+) 4.72 kN/mm (-) 7.08 kN/mm Fy  CUREE (+) 39.31 kN (-) -41.75 kN ∆y  CUREE (+) 8.32 mm (-) -5.90 mm Figure  F.2.10    Cyclic test result for test combination CLT-A-N-P-C04 205    MECHANICAL PROPERTIES Fmax (+) 51.98 kN (-) -54.83 kN dFmax  (+) 40.23 mm (-) -23.19 mm 0.1 Fmax (+) 5.20 kN (-) -5.48 kN 0.4 Fmax (+) 20.79 kN (-) -21.93 kN 0.8 Fmax (+) 41.58 kN (-) -43.86 kN D (+) 6.12 [-] (-) 4.00 [-] Ke (+) 5.54 kN/mm (-) 5.98 kN/mm Fy  CUREE (+) 45.49 kN (-) -47.24 kN ∆y  CUREE (+) 8.21 mm (-) -7.90 mm Figure  F.2.11    Cyclic test result for test combination CLT-A-N-P-C05 206    MECHANICAL PROPERTIES Fmax (+) 52.39 kN (-) -54.15 kN dFmax  (+) 32.37 mm (-) -18.02 mm 0.1 Fmax (+) 5.24 kN (-) -5.41 kN 0.4 Fmax (+) 20.96 kN (-) -21.66 kN 0.8 Fmax (+) 41.91 kN (-) -43.32 kN D (+) 4.59 [-] (-) 3.68 [-] Ke (+) 5.16 kN/mm (-) 6.14 kN/mm Fy  CUREE (+) 45.76 kN (-) -51.25 kN ∆y  CUREE (+) 8.87 mm (-) -8.35 mm Figure  F.2.12    Cyclic test result for test combination CLT-A-N-P-C06 207    MECHANICAL PROPERTIES Fmax (+) 51.10 kN (-) -54.56 kN dFmax  (+) 32.81 mm (-) -17.43 mm 0.1 Fmax (+) 5.11 kN (-) -5.46 kN 0.4 Fmax (+) 20.44 kN (-) -21.83 kN 0.8 Fmax (+) 40.88 kN (-) -43.65 kN D (+) 3.70 [-] (-) 5.61 [-] Ke (+) 4.12 kN/mm (-) 7.16 kN/mm Fy  CUREE (+) 45.44 kN (-) -50.48 kN ∆y  CUREE (+) 11.04 mm (-) -7.05 mm Figure  F.2.13    Cyclic test result for test combination CLT-A-N-P-C07 208 F.3 Pictures  Figure  F.3.1    Front view of Bracket A with 18 Spiral Nails under monotonic loading, parallel to the grain  Figure  F.3.2    Pull-out failure of Spiral Nails in combination with Bracket A  209 Figure  F.3.3    Failed connection with Bracket A; Pull-out failure under monotonic loading  Figure  F.3.4    CLT panel after cyclic loading perpendicular to the grain; combined pull-out failure with shear failure of fasteners  210 Figure  F.3.5    Stages 1 to 4 of Bracket A with Spiral Nails under cyclic loading  211 Figure  F.3.6    Stages 5 to 8 of Bracket A with Spiral Nails under cyclic loading  212 Figure  F.3.7    Bracket A with 18 Spiral Nails under monotonic loading perpendicular to the grain  213Appendix G  Tube with diameter 50.8 mm (2 inch) G.1 Monotonic Test Results  Test ID   T2-1-M04 T2-1-M06 T2-1-M07 Average Fmax [kN] 59.8 62.3 51.8 59.6 dFmax [mm] 20.0 25.0 10.0 20.0 Fult [kN] 47.8 49.8 41.5 47.7 ∆ult [mm] 21.0 26.0 11.0 23.1 D  [-]       6.2 Ke [kN/mm]       14.1 Fy - CUREE [kN]       52.3 ∆y - CUREE [mm]       3.7 Figure  G.1.1    Monotonic test results for connection combination T2-1-M  214G.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 60.73 kN (-) - kN dFmax  (+) 14.89 mm (-) - mm 0.1 Fmax (+) 6.07 kN (-) - kN 0.4 Fmax (+) 24.29 kN (-) - kN 0.8 Fmax (+) 48.59 kN (-) - kN D (+) 8.41 [-] (-) - [-] Ke (+) 26.36 kN/mm (-) - kN/mm Fy  CUREE (+) 57.19 kN (-) - kN ∆y  CUREE (+) 2.17 mm (-) - mm Figure  G.2.1    Cyclic test result for test combination CLT-T2-1-C01 215    MECHANICAL PROPERTIES Fmax (+) 53.94 kN (-) - kN dFmax  (+) 15.20 mm (-) - mm 0.1 Fmax (+) 5.39 kN (-) - kN 0.4 Fmax (+) 21.58 kN (-) - kN 0.8 Fmax (+) 43.15 kN (-) - kN D (+) 14.68 [-] (-) - [-] Ke (+) 43.65 kN/mm (-) - kN/mm Fy  CUREE (+) 52.38 kN (-) - kN ∆y  CUREE (+) 1.20 mm (-) - mm Figure  G.2.2    Cyclic test result for test combination CLT-T2-1-C02 216    MECHANICAL PROPERTIES Fmax (+) 57.45 kN (-) - kN dFmax  (+) 15.34 mm (-) - mm 0.1 Fmax (+) 5.74 kN (-) - kN 0.4 Fmax (+) 22.98 kN (-) - kN 0.8 Fmax (+) 45.96 kN (-) - kN D (+) 4.64 [-] (-) - [-] Ke (+) 14.37 kN/mm (-) - kN/mm Fy  CUREE (+) 55.52 kN (-) - kN ∆y  CUREE (+) 3.86 mm (-) - mm Figure  G.2.3    Cyclic test result for test combination CLT-T2-1-C03 217    MECHANICAL PROPERTIES Fmax (+) 53.27 kN (-) - kN dFmax  (+) 14.22 mm (-) - mm 0.1 Fmax (+) 5.33 kN (-) - kN 0.4 Fmax (+) 21.31 kN (-) - kN 0.8 Fmax (+) 42.62 kN (-) - kN D (+) 7.79 [-] (-) - [-] Ke (+) 19.72 kN/mm (-) - kN/mm Fy  CUREE (+) 49.77 kN (-) - kN ∆y  CUREE (+) 2.52 mm (-) - mm Figure  G.2.4    Cyclic test result for test combination CLT-T2-1-C04 218    MECHANICAL PROPERTIES Fmax (+) 53.54 kN (-) - kN dFmax  (+) 6.21 mm (-) - mm 0.1 Fmax (+) 5.35 kN (-) - kN 0.4 Fmax (+) 21.42 kN (-) - kN 0.8 Fmax (+) 42.83 kN (-) - kN D (+) 3.60 [-] (-) - [-] Ke (+) 15.07 kN/mm (-) - kN/mm Fy  CUREE (+) 50.84 kN (-) - kN ∆y  CUREE (+) 3.37 mm (-) - mm Figure  G.2.5    Cyclic test result for test combination CLT-T2-1-C05 219 G.3 Pictures  Figure  G.3.1    Test set-up of T2  Figure  G.3.2    Loading stages of T2; no visible destruction of CLT panel  220 Figure  G.3.3    Destroyed Tube T2; CLT panel without any failure  221Appendix H  Tube with diameter 76.2 mm (3 inch) H.1 Monotonic Test Results   Test ID   T3-1-M01 T3-1-M02 T3-1-M03 T3-1-M04 Average Fmax [kN] 48.7 51.7 51.4 47.7 49.3 dFmax [mm] 25.0 24.0 26.0 21.0 24.0 Fult [kN] 39.0 41.4 41.1 38.1 39.4 ∆ult [mm] 35.0 35.0 33.0 26.0 29.6 D  [-]         10.1 Ke [kN/mm]         14.1 Fy - CUREE [kN]         41.2 ∆y - CUREE [mm]         2.9 Figure  H.1.1    Monotonic test results for connection combination T3-1-M  222H.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 34.19 kN (-) - kN dFmax  (+) 10.81 mm (-) - mm 0.1 Fmax (+) 3.42 kN (-) - kN 0.4 Fmax (+) 13.68 kN (-) - kN 0.8 Fmax (+) 27.35 kN (-) - kN D (+) 9.20 [-] (-) - [-] Ke (+) 17.78 kN/mm (-) - kN/mm Fy  CUREE (+) 32.40 kN (-) - kN ∆y  CUREE (+) 1.82 mm (-) - mm Figure  H.2.1    Cyclic test result for test combination CLT-T3-1-C01 223    MECHANICAL PROPERTIES Fmax (+) 32.67 kN (-) - kN dFmax  (+) 11.48 mm (-) - mm 0.1 Fmax (+) 3.27 kN (-) - kN 0.4 Fmax (+) 13.07 kN (-) - kN 0.8 Fmax (+) 26.14 kN (-) - kN D (+) 6.12 [-] (-) - [-] Ke (+) 10.46 kN/mm (-) - kN/mm Fy  CUREE (+) 30.76 kN (-) - kN ∆y  CUREE (+) 2.94 mm (-) - mm Figure  H.2.2    Cyclic test result for test combination CLT-T3-1-C02 224    MECHANICAL PROPERTIES Fmax (+) 36.36 kN (-) - kN dFmax  (+) 11.20 mm (-) - mm 0.1 Fmax (+) 3.64 kN (-) - kN 0.4 Fmax (+) 14.55 kN (-) - kN 0.8 Fmax (+) 29.09 kN (-) - kN D (+) 14.25 [-] (-) - [-] Ke (+) 29.71 kN/mm (-) - kN/mm Fy  CUREE (+) 32.00 kN (-) - kN ∆y  CUREE (+) 1.08 mm (-) - mm Figure  H.2.3    Cyclic test result for test combination CLT-T3-1-C03 225    MECHANICAL PROPERTIES Fmax (+) 35.59 kN (-) - kN dFmax  (+) 15.59 mm (-) - mm 0.1 Fmax (+) 3.56 kN (-) - kN 0.4 Fmax (+) 14.24 kN (-) - kN 0.8 Fmax (+) 28.47 kN (-) - kN D (+) 9.06 [-] (-) - [-] Ke (+) 16.71 kN/mm (-) - kN/mm Fy  CUREE (+) 33.02 kN (-) - kN ∆y  CUREE (+) 1.98 mm (-) - mm Figure  H.2.4    Cyclic test result for test combination CLT-T3-1-C04 226    MECHANICAL PROPERTIES Fmax (+) 47.41 kN (-) - kN dFmax  (+) 4.95 mm (-) - mm 0.1 Fmax (+) 4.74 kN (-) - kN 0.4 Fmax (+) 18.97 kN (-) - kN 0.8 Fmax (+) 37.93 kN (-) - kN D (+) 7.84 [-] (-) - [-] Ke (+) 16.18 kN/mm (-) - kN/mm Fy  CUREE (+) 43.47 kN (-) - kN ∆y  CUREE (+) 2.69 mm (-) - mm Figure  H.2.5    Cyclic test result for test combination CLT-T3-1-C05 227    MECHANICAL PROPERTIES Fmax (+) 38.08 kN (-) - kN dFmax  (+) 19.77 mm (-) - mm 0.1 Fmax (+) 3.81 kN (-) - kN 0.4 Fmax (+) 15.23 kN (-) - kN 0.8 Fmax (+) 30.47 kN (-) - kN D (+) 6.19 [-] (-) - [-] Ke (+) 9.89 kN/mm (-) - kN/mm Fy  CUREE (+) 36.03 kN (-) - kN ∆y  CUREE (+) 3.64 mm (-) - mm Figure  H.2.6    Cyclic test result for test combination CLT-T3-1-C06 228 H.3 Pictures  Figure  H.3.1    Test set-up for T3 under cyclic loading  Figure  H.3.2    Deformed tube T3 under monotonic loading  229 Figure  H.3.3    Close up of failed tube T3  230 Figure  H.3.4    Deformed tube T3 under cyclic loading  231Appendix I  Tube with diameter 101.6 mm (4 inch) I.1 Monotonic Test Results   Test ID   T4-1-M01 T4-1-M02 T4-1-M03 T4-1-M04 Average Fmax [kN] 42.7 42.9 40.2 47.3 41.8 dFmax [mm] 18.0 18.0 19.0 17.0 17.0 Fult [kN] 34.1 34.3 32.2 37.9 33.4 ∆ult [mm] 21.0 21.0 20.0 18.0 19.9 D  [-]         3.1 Ke [kN/mm]         5.3 Fy - CUREE [kN]         33.9 ∆y - CUREE [mm]         6.4 Figure  I.1.1    Monotonic test results for connection combination T4-1-M  232I.2 Cyclic Test Results    MECHANICAL PROPERTIES Fmax (+) 36.26 kN (-) - kN dFmax  (+) 12.93 mm (-) - mm 0.1 Fmax (+) 3.63 kN (-) - kN 0.4 Fmax (+) 14.50 kN (-) - kN 0.8 Fmax (+) 29.00 kN (-) - kN D (+) 5.94 [-] (-) - [-] Ke (+) 11.63 kN/mm (-) - kN/mm Fy  CUREE (+) 30.82 kN (-) - kN ∆y  CUREE (+) 2.65 mm (-) - mm Figure  I.2.1    Cyclic test result for test combination CLT-T4-1-C05 233    MECHANICAL PROPERTIES Fmax (+) 36.38 kN (-) - kN dFmax  (+) 14.04 mm (-) - mm 0.1 Fmax (+) 3.64 kN (-) - kN 0.4 Fmax (+) 14.55 kN (-) - kN 0.8 Fmax (+) 29.10 kN (-) - kN D (+) 2.57 [-] (-) - [-] Ke (+) 5.59 kN/mm (-) - kN/mm Fy  CUREE (+) 31.32 kN (-) - kN ∆y  CUREE (+) 5.60 mm (-) - mm Figure  I.2.2    Cyclic test result for test combination CLT-T4-1-C02 234    MECHANICAL PROPERTIES Fmax (+) 39.01 kN (-) - kN dFmax  (+) 18.28 mm (-) - mm 0.1 Fmax (+) 3.90 kN (-) - kN 0.4 Fmax (+) 15.60 kN (-) - kN 0.8 Fmax (+) 31.21 kN (-) - kN D (+) 3.91 [-] (-) - [-] Ke (+) 6.62 kN/mm (-) - kN/mm Fy  CUREE (+) 33.51 kN (-) - kN ∆y  CUREE (+) 5.06 mm (-) - mm Figure  I.2.3    Cyclic test result for test combination CLT-T4-1-C03 235    MECHANICAL PROPERTIES Fmax (+) 34.90 kN (-) - kN dFmax  (+) 14.89 mm (-) - mm 0.1 Fmax (+) 3.49 kN (-) - kN 0.4 Fmax (+) 13.96 kN (-) - kN 0.8 Fmax (+) 27.92 kN (-) - kN D (+) 3.59 [-] (-) - [-] Ke (+) 6.51 kN/mm (-) - kN/mm Fy  CUREE (+) 29.64 kN (-) - kN ∆y  CUREE (+) 4.55 mm (-) - mm Figure  I.2.4    Cyclic test result for test combination CLT-T4-1-C04 236    MECHANICAL PROPERTIES Fmax (+) 35.52 kN (-) - kN dFmax  (+) 18.15 mm (-) - mm 0.1 Fmax (+) 3.55 kN (-) - kN 0.4 Fmax (+) 14.21 kN (-) - kN 0.8 Fmax (+) 28.42 kN (-) - kN D (+) 2.91 [-] (-) - [-] Ke (+) 5.06 kN/mm (-) - kN/mm Fy  CUREE (+) 32.55 kN (-) - kN ∆y  CUREE (+) 6.43 mm (-) - mm Figure  I.2.5    Cyclic test result for test combination CLT-T4-1-C05  237I.3 Pictures  Figure  I.3.1    Test set-up tube T4  Figure  I.3.2    Deformed tube T4 after cyclic loading  238  Figure  I.3.3    Close-up of tested tube T4  239 Figure  I.3.4    MTS test machine set-up for tube-type connections with base steel beam  240  Figure  I.3.5    Tubes T2, T3, and T4 after loaded with cyclic loading protocol  241Appendix J  Damage assessment of cross-laminated timber connections and shear wall connections subjected to simulated earthquake loads  J.1 Introduction In Chapter 2, a series of CLT bracket connections was tested, analyzed, applied to an accumulative damage model, and discussed. Based on the observed results, a damage scale was developed. Here, the obtained connection results were expanded on entire shear wall tests, which were conducted at FPInnovations. FPInnovations had started several testing series in 2009 to investigate the material properties of CLT. The conducted research focused on shear walls and their connection to the foundation (Schneider, 2009). Another research carried out focused on connections between wall and deck, as well as wall to wall under monotonic and seismic loadings. Using the data from individual connection and shear wall tests, in this chapter, Kratzig’s damage accumulation principles are applied to both test series (Williams and Sexsmith, 1995). The calculated damage stages will be compared with visual observations. J.2 Connection test The connection tests were conducted separately for loading parallel-to-the-grain and perpendicular-to-the-grain. In Section 2.2.1 the connection testing materials and procedures are described in detail. J.3 Shear wall test All shear walls were tested on the same test set-up with minor changes for each different test. Each wall panel sat on a solid base beam, which provided holes for all wall set-ups. The wall test set-up is schematically illustrated in Figure  J.4.1. The wall length varied between 2.3 m and 3.45 m. The brackets were connected to the base beam with bolts and washers either 12.7 mm (1/2") or  9.52 mm (3/8") depending on the bracket used, and connected with different fasteners to the CLT wall panel. The vertical load was applied to the wall panel by a top beam that was attached to the top edge side of the wall panel. The vertical load applied was 20 kN/m. Depending on the length, either two or four vertical actuators equipped with a load  242cell applied a constant force between 23 kN and 69 kN. To allow maximum freedom of rotation of the panel, the horizontal force was induced through a pinned connection on each end of the wall. In the starting position of the test, the centre of the actuator lines up with the top edge of the wall. The main actuator was connected to a tall braced tower. To prevent tipping out of plane of the wall panel, there were two rollers on either side of the top beam to provide lateral support at the top end of the wall. An overview of the tested walls which will be considered for the damage accumulation indices is provided in Table  J.4.1. J.4 Loading protocol for shear wall tests For the shear wall tests, the same type loading protocols as for the connection tests were used. The monotonic test was loaded horizontally unidirectional and in-plane at a loading rate of 10.16 mm (0.4") per minute. The displacement-controlled loading stopped at 60% of the peak load after the peak load was reached. Using the obtained data, the displacement at 80% of the peak load is calculated which is necessary for the cyclic tests. For the cyclic CUREE protocol, a displacement rate of 5.08 mm/s (0.2"/s) was chosen. A schematic plot time vs. displacement of the cyclic loading is shown in Figure 2.10.    243 Figure  J.4.1    Isometric drawing of the shear wall test set-up  244Table  J.4.1    Shear wall tests used for analysis  Wall Configuration Test No. Connections Loading Protocol 1  00C 4 Bracket A SN 16d, n=18 CUREE 01M Ramp 01C CUREE 03 CUREE 04 RN 10d, n=12 CUREE 05 S1 n=18 CUREE 06 S2 n=9 CUREE  20 7 Bracket A SN 16d, n=18 CUREE 3  09M 4 Bracket B SN 16d, n=10 Ramp 09C CUREE 10 CUREE 5  13M 9 Bracket B SN 16d, n=10 Ramp 13C CUREE 6  15M 9 Bracket B SN 16d, n=10 in step joint  SFS 1, n=8 Ramp 15C CUREE 19 7 Bracket B SN 16d, n=10 CUREE SN = Spiral nail 3.9 × 89mm, RN = Ring shank nail 2.4 × 76mm, S1 = Screw 4 × 70mm, S2 = Screw 5 × 90mm, TR65 = Timber rivet 65mm, TR90 = Timber rivet 90mm.   245J.5 Experimental results and calibration of damage indices  J.6 Failure modes of shear wall connections In Section 0 various failure modes in CLT connections were discussed. The observed failure modes in CLT shear walls connected to a steel base beam were similar to the connection tests (Figure  J.7.1). Pull-out failure of the fasteners dominated for tests with nails. Screws often showed wood crushing or fasteners failed in shear. The end brackets connected with 5 x 90 mm screws (Wall 06) failed with bracket rupture. In some cases the CLT panel experienced block shear failure. Edge breakout of the panel occurred often in combination with the pull-out failure of the fasteners. Regardless of the sizes of the wall panels, all failure modes were initiated but rocking of the panel, which results in an up and down movement, thus parallel-to-the–grain loading on the connection. The horizontal sliding of the panel was neglegtably small (3mm to 4mm), and therefore, the failures, which were found in the perpendicular-to-the-grain direction in the connection tests, did not influence the failure modes  J.7 Shear wall analysis The shear wall tests were divided into eight groups based on size and connections. For this research, four relevant groups are considered and shown in Table  J.4.1. The damage assessment was carried out based on the cyclic data sets for 16 walls. The damage-time curves of a wall with the size of 2.3m × 2.3m and various connectors are compared in Figure  J.7.2 to Figure  J.7.5. The validation of the calculated value was done by visual observation of the test walls. The first group of wall tests shows a similar behaviour to the observed in the connection tests. The curves for walls with spiral nails can be found at the lower end of the curves (Figure  J.7.2; curve 00, 01C, 03). It is interesting to note that in group 3, both tests performed identical.  246 a) Fastener withdrawal  b) Shear fracture of fastener   c) Edge break out at the underside of the sample  d) Wood crushing  e) CLT blockshear failure  f) Bracket fracture Figure  J.7.1    Failure modes observed in CLT shear wall tests  247  Figure  J.7.2    Wall results for group 1 (2.3m x 2.3)   248 Figure  J.7.3    Wall results for group 3 (2.3m x2.3m)  249 Figure  J.7.4    Wall results for group 5 (2.3m x 3.45m)  250 Figure  J.7.5    Wall results for group 6 (3.45m x 2.3m)  The results from Group 6 show the influence of the wall panel on the damage accumulation (Figure  J.7.5). Wall #19 consists of 3 separate panels connected with screws in the step joints. Even with a higher number of brackets, the slope of the curve is very low. Wall 15C consisting of one panel with 2 openings and loaded with the same protocol, shows significantly higher rates of the damage accumulation. It shows the necessity of having the same boundary conditions to be able to compare the damage results.  251J.8 Damage prediction  In this section, Krätzig’s index (energy-based index) was applied to full size CLT shear walls. In order to make the proposed damage model useful for predicting and evaluating the damage state of CLT connections, the relationship between calculated damage index and observed damage needs to be established. To address the observed damage in the conducted tests, damage states for CLT connections in shearwalls need to be defined accordingly to the connection tests. Previous research considered the final stage after completing the loading protocol for the damage indices. The focus for CLT connections was placed on the development of the damage accumulation over time. The results have to be validated with the destructivity of the connection. The limit states from the connection tests were determined as None, Minor, Moderate, Severe, and Collapse. The relationship between the damage index and the observed damage is shown in Table  J.8.1 and has to be validated with shear wall tests.  Table  J.8.1    Classification of Damage for CLT connections Degree of damage  Damage description Damage index scale None  No visible damage observed D < 0.20 Minor   Minor pull-out of fasteners; light plastic deformation of  bracket; minor repairs are required 0.20 ≤ D < 0.35 Moderate   Visual permanent deflections of bracket; shear failure of  small number of fasteners; extensive pull-out of fasteners;  0.35 ≤ D < 0.65 Severe      Major or complete failure of fasteners; severe crack in  bracket; separation of bracket from CLT panel; requires  replacement of bracket in different position at CLT wall  to be serviceable again; severe wood crushing in outer  layer of CLT 0.65 ≤ D < 0.75 Collapse  Total or partial collapse of connection D > 0.75  

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