10 The sliding hinge joint (SHJ) is a type of suppleme ntal energy dissipation system for column 11 bases or beam-to-column connections of steel Moment Resisting Frames (MRFs). It is based 12 on the application of symmetric/asymmetric friction dampers in joints to develop a 13 dissipative mechanism alternative to the column/bea m yielding. This typology was initially 14 proposed in New Zealand and, more recently, is star ting to be tested and applied also in 15 Europe. While on the one hand this technology provi des great benefits such as the damage 16 avoidance, on the other hand, due to the high unloa ding stiffness of the dampers in tension or 17 compression, its cyclic response is typically chara cterized by a limited self-centering capacity. 18 To address this shortcoming, the objective of the w ork herein presented is to examine the 19 possibility to add to these connections also a self -centering capacity proposing new layouts 20 based on a combination of friction devices (providi ng energy dissipation capacity), pre-loaded 21 threaded bars and disk springs (introducing in the joint restoring forces). 22 In this paper, as a part of an ongoing wider experi mental activity regarding the behaviour of 23 self-centering connections, the attention is focuse d on the problem of achieving the self- 24 centering of the column bases of MRFs by studying a detail consisting in a column-splice 25 equipped with friction dampers and threaded bars wi th Belleville disk springs, located above 26 a traditional full-strength column base joint. The main benefits obtained with the proposed 27 layout are that: i) the self-centering capability i s obtained with elements (threaded bars and 28 Belleville springs) which have a size comparable to the overall size of the column-splice cover 29 plates; ii) all the re-centering elements are moved far from the concrete foundation avoiding 30 any interaction with the footing. The work reports the main results of an experimental 31 investigation and the analysis of a MRF equipped wi th the proposed column base joints. 32 33 Keywords: Friction dampers, self-centering, FREEDAM , base-plate, modelling 34 -2- 1. INTRODUCTION 35 36 Eurocodes require to design structures to ensure th e achievement of minimum performance 37 levels under a set of design load combinations [1,2 ]. Current design procedures are based on 38 structural checks for Serviceability Limit States ( SLS) (related to the most frequent conditions 39 occurring during the life-time of the structure) an d for Ultimate Limit States (ULS) for which 40 the structure, in case of rare seismic events, can be designed to dissipate energy in selected 41 zones. 42 The modern seismic protection strategies implemente d into international building codes are 43 based, in case of destructive seismic events, on th e absorption of the seismic energy in 44 dissipative zones, which are detailed to sustain cy clic inelastic rotation demands [3]. In case of 45 steel Moment Resisting Frames (MRFs) this strategy is traditionally applied by properly over- 46 strengthening columns and connections enforcing, in this manner, the development of plastic 47 hinges in the beam ends and at the base of the colu mns. Additionally, to maximize the energy 48 dissipation, the plastic zones are spread along the elevation of the building, promoting the 49 development of a global failure mode through the ap plication of members hierarchy criteria 50 and the design of full strength connections [4-7]. Therefore, owing to the assumptions made 51 in design, traditional procedures typically lead to structures characterized by weak beams 52 and column bases, with strong joints. 53 This approach, if on the one hand provides benefits , such as the development of a stable 54 plasticization and the reduction of the inter-store y drifts under serviceability loading 55 conditions, on the other hand, leads to significant shortcomings. The most substantial 56 weakness is intrinsic in the design strategy itself . In fact, although the damage is needed to 57 absorb the input earthquake energy, it also represe nts one of the main sources of economic 58 loss [8-11]. In fact, since the dissipative zones a re constituted by sections or elements 59 belonging to the structural system, after severe se ismic events, the structure is affected by 60 significant damage and, because of permanent plasti c deformations, it is characterized by a 61 pattern of residual drifts. In general, the magnitu de of this out-of-plumbness may be 62 significant in view of the actual possibility to re pair the structure after a destructive seismic 63 event. 64 Aiming to design structures undergoing minimal dama ge, special typologies of dissipative 65 partial strength joints based on the inclusion of f riction dampers in connections have been 66 proposed and, recently, extensive studies have been carried out in research programs 67 -3- worldwide [12-15]. These connections were initially proposed by Grigorian and co-authors in 68 1993 [16] and, subsequently, many other theoretical , experimental and modelling works, as 69 well as practical applications, were carried out, e specially in New Zealand, developing the so- 70 called called Sliding Hinge Joint (SHJ). This conne ction is characterised by very simple details 71 based on the inclusion of Asymmetric Friction Conne ctions (AFC) or Symmetric Friction 72 Connections (SFC) at the bottom beam flange, with f riction pads made of mild steel, 73 aluminium, brass or – in the most recent versions – abrasion-resistant steel (e.g. [16-22]). 74 Similar solutions were also patented in 2000 in Jap an [23,24] while, more recently, other 75 alternatives have been proposed suggesting new layo uts, in which the friction damper is 76 conceived as a separate element fabricated in the s hop and fastened on site to the beam 77 bottom flange [25-28]. This layout, which is probab ly not as simple as the SHJ, provides the 78 possibility to realise the whole damper in the shop allowing a better control on the materials 79 quality (e.g. higher control of the surface conditi ons, continuous factory controls on the 80 production, control on the employed bolts quality), and on the application of rigorous bolts 81 installation procedures complying with the relevant European standards [28-34]. The layout 82 of the typical beam-to-column joint, recently propo sed in Europe for application in semi- 83 continuous steel Moment Resisting Frames (MRFs), re presents an alternative to a stiffened 84 Double Split Tee connection (DST) where, in place o f the bottom Tee, a slotted friction device 85 with a haunch slipping on friction shims pre-stress ed with pre-loadable high strength bolts 86 (Fig.1) is realised. All the elements of the connec tion constitute a Symmetrical Friction 87 Connection (SFC) which is, as already underlined, a friction damping device fabricated as a 88 standalone element in the shop. With such detail th e beam is forced to rotate around the pin 89 located at the base of the upper T-stub web and the energy dissipation is ensured by the 90 alternate slippage of the lower beam flange on fric tion shims (Fig.1). 91 -4- 92 Fig. 1 – Typical layout of one of the connections studied in [28] 93 94 This connection, similarly to the SHJ, should be im plemented to behave rigidly at SLSs and to 95 allow the beam-to-column inelastic rotation at the ULS. Additionally, through the application 96 of proper hierarchy criteria, both at the global an d local level, it can be easily designed to be 97 the only source of energy dissipation of the whole structure. 98 Within this framework, considering the encouraging outcomes of previous research projects 99 dealing with the application of such connections, i n this paper, the problem of the self- 100 centering structures equipped with dissipative fric tion joints is analysed. In fact, due to 101 permanent deformations in the friction dampers, sim ilarly to what occurs when plastic zones 102 are concentrated in the beams or in yielding connec tions, significant out-of-plumbness 103 displacements can remain after a severe ground moti on [15, 42-44]. Indeed, although these 104 connections are very effective from the point of vi ew of the damage avoidance, they still 105 provide significant problems related to the low sel f-centering capacity. This drawback is 106 mainly due to the high unloading stiffness of the f riction dampers in tension or compression. 107 To avoid this undesired behaviour, as already propo sed in several past studies [34-41] a 108 supplemental re-centering system can be adopted. 109 Specifically, in this paper, the attention is focus ed on the problem of self-centering the column 110 base joint, by studying a detail consisting in a co lumn-splice equipped with friction dampers 111 and threaded bars with Belleville disk springs, loc ated just above a traditional full-strength 112 base plate joint. The main advantages of the propos ed layout are that: i) the self-centering 113 capability is obtained with re-centering elements ( threaded bars and Belleville springs) which 114 -5- have a small size, similar to the dimension of the column-splice cover plates; ii) all the re- 115 centering elements are moved far from the concrete foundation. The work reports the main 116 results of an experimental investigation and prelim inary analyses of MRFs equipped with re- 117 centering FREEDAM column base joints. The obtained results are hereinafter critically 118 discussed showing the promising performances of the proposed column base connection. 119 120 2. CONCEPT OF THE PROPOSED CONNECTION 121 2.1 Friction dampers and re-centering systems 122 Being an effective way of dissipating energy, dampe rs based on principles of dry friction have 123 become very popular and are largely used in high ri sk seismic zones. In the last decades, the 124 application of this concept has been subject of num erous studies [35-38, 45] and many 125 friction dampers have been proposed for practical p urposes. This damper typology usually 126 dissipates energy through the alternate slippage of at least two surfaces in contact, on which a 127 transversal clamping force is applied with hydrauli c systems [46], electromagnetic forces [47] 128 or, in the simplest case, by means of mechanical de vices such as high strength bolts. This last 129 clamping method is the most common in civil enginee ring practice. 130 The cyclic behaviour of friction dampers is normall y characterized by a rigid-plastic 131 hysteresis which depends only on two parameters: th e clamping force and the friction 132 coefficient of the interfaces in contact. The first parameter is usually governed by the 133 application tightening procedures which are based e ssentially on the control of the nut 134 rotation (displacement-controlled procedure), appli ed torque (force-controlled procedure) or 135 on the employment of specific devices which fail or squash at the achievement of the proof 136 preloading level (e.g. DTI or squirter DTI washers and HRC bolts) [26]. Conversely, the second 137 parameter (namely the friction coefficient) is pred icted by means of physical modelling or 138 experimental testing. In the former case, the atten tion is focused on the modelling of complex 139 and microscopic phenomena such as adhesion and plou ghing which are dependent upon the 140 surfaces topography, the materials hardness, the me chanical properties and the effects of 141 interface layers. In the latter case, which is the most common in structural engineering 142 practice, conversely, the properties of the frictio n interface are studied by means of 143 experimental testing which, for seismic engineering purposes is generally considered 144 sufficient to provide the information needed for de signing the devices. A general discussion 145 dealing with the main factors influencing the frict ion interaction is reported in [10,48,49]. 146 -6- The main proposals of application of friction dampe rs in steel structures are referred to 147 bracing systems or beam-to-column connections. One of the first devices based on friction 148 was that developed in [50] which introduced at the intersection of braces, brake lining pads 149 between the steel sliding surfaces. One of the simp lest forms of friction damper has been 150 proposed in [51] who adopted simple bolted slotted plates located at the end of a 151 conventional bracing member. The brace-to-frame con nection was designed to slip with fully 152 elastic braces. Another friction damper for chevron braces was proposed in [52]. Concerning 153 connections, as previously said, application of pri nciples of dry friction were initially first 154 developed by Grigorian and co-authors [16] and subs equently extensive studies were carried 155 out in New Zealand by the research group at the Uni versity of Auckland [9,10, 17-22, 53,54] 156 and in other countries applying these principles al so to other structural typologies [55,56]. 157 More recently, other works on a specific type of sl iding hinge joint have been performed also 158 in Europe in a research activity regarding the anal ysis of friction materials, the bolt 159 installation procedures, the long-term response due to relaxation of the slip force, the 160 robustness assessment, the FE modelling and the exp erimental analysis of real-scale 161 structures or sub-assemblies of joints [26,28]. 162 While, as herein summarized, friction dampers in be am-to-column connections have been 163 largely investigated, the application of friction d ampers in column base joints of steel 164 structures is only a recent proposal and little kno wledge is currently available. The idea to 165 dissipate energy in the base plate with friction de vices comes from the field observation of 166 damages after the earthquakes of Northridge (1994), Kobe (1995) and Tohuku (2011). In fact, 167 during the technical surveys, in many cases, severe damage involving plate and anchor bolts 168 was observed. Additionally, past experimental tests have indicated that the traditional base 169 plate connections are prone to the development of d amage into elements which are not easy 170 to replace such as the base section of the column ( in case of full-strength connections) or the 171 base plate/anchors (in case of partial strength con nections) and, due to residual 172 deformations, may give rise to a pattern of residua l lateral displacements in the whole 173 building. Therefore, in general, owing to the limit ed dissipative capacity and difficult 174 reparability of the base joint (they are typically hidden by the flooring of the first storey) the 175 occurrence of damage at the base of the building re presents a significant shortcoming both in 176 view of the actual reparability of the building and in terms of economic cost to be sustained in 177 the aftermath. All these issues have recently motiv ated a significant number of research 178 activities worldwide dealing with the development o f innovative base plate connections 179 -7- equipped with dampers able to limit damage, while p reserving the ability of the structure to 180 dissipate energy in case of rare seismic events. Th ese connections, in some cases, have been 181 equipped also with re-centering elements able to re store the columns to the initial position. 182 Two layouts were proposed by McRae and co-authors [ 58], while in [59] the study of the 183 efficiency of the dissipation of seismic energy thr ough column base solutions has been 184 performed carrying out a series of experimental tes ts on different low damage steel base 185 connection. Within this work, two new design soluti ons were tested: the weak axis aligned 186 asymmetric friction connection, where friction surf aces are parallel to the web on plates 187 outstanding from the column flange, the strong axis aligned asymmetric friction connection, 188 where friction surfaces are parallel to the column flange. It is worth noting that, as evidenced 189 in [58], critical phenomena occurring with conventi onal full-strength connections can be 190 mitigated by means of friction column base connecti on, such as that proposed in this paper. In 191 fact, while in traditional frames due to yielding o f the column base section and local buckling 192 phenomena, axial shortening of the column may occur [58, 60], with damage free connections, 193 such as the double friction base columns suggested in [58, 59] owing to the absence of plastic 194 deformations in the column, the axial shortening an d its detrimental effects can be completely 195 avoided. Recently a novel type of rocking damage fr ee connection has been proposed in [15]. 196 This column base, has a circular hollow section wel ded to a thick plate, four post-tensioning 197 tendons to give a self-centering capacity to the co nnection and friction dampers to dissipate 198 energy. 199 Other practical cases of self-centering systems pro posed in literature usually include a tendon, 200 applied in the joint or over the entire extension o f the structure. In [39] it was proposed to 201 include friction ring springs to the SHJ, obtaining a flag-shape behavior of the connection. A 202 similar approach in terms of re-centering was propo sed in [40] who employed as re-centering 203 component rods applied at the tips of the whole bea m, rather than only on the joint. In this 204 work, the introduction of an “active link” was sugg ested and the connection to the beam, at 205 both ends, was achieved by means of pre-tensioned r ods. In the second work, the employment 206 of a set of rods going through the entire segment o f the beam and attached in the joint section 207 was proposed. 208 Self-centering base connections have also been deve loped in [41] using post-tensioned rods 209 anchored to the column foundation. The aim is to en sure the possibility of movement, pre- 210 stressing the rods within their elastic capacity. H owever, the proposed solutions, based on 211 anchoring the rods to the foundation, can be less e ffective in a replacement situation. Friction 212 -8- systems can also show a self-centering ability when employed with an asymmetric 213 configuration of the damper [21]. However, such cap ability is usually limited, and additional 214 components are normally needed to restore the conne ction itself or the structure. A 215 significant practical implementation of the damage avoidance design strategy is described in 216 [57]. In this project, the building is designed in the transverse direction with tension limited 217 rocking shear walls and in the opposite direction w ith Sliding Hinge Joint MRFs. In this 218 application, the rocking shear walls are equipped w ith Ringfederer springs to obtain the self- 219 centering ensuring hinge formation under a stable r ocking mechanism. Conversely, the MRF 220 bays are equipped with conventional SHJs without se lf-centering devices. The similarities 221 between the solutions adopted in [57] and the appli cation described in this paper are related 222 to the adoption of heavy-load springs to adjust the capacity of the structure and the 223 introduction of friction dampers in the column base . Nevertheless, as a difference, the 224 connection hereinafter presented proposes to introd uce in the column base a simple system 225 of threaded bars with sets of Belleville washers ac ting as a spring to provide the needed self- 226 centering action. This proposal wants to keep the l ayout of the connection as simple as 227 possible providing, other than the self-centering c apacity, additional benefits such as the 228 absence of interaction with the concrete foundation and the limited size of the connection 229 which is, overall, similar or lower than the size o f the cover plates employed to realize a 230 traditional column splice connection. 231 232 2.2 Proposed Solution 233 The proposed connection consists in a slotted colum n splice equipped with friction pads 234 located above a traditional full-strength base plat e joint (Fig. 2a) [38]. In particular, 235 symmetrical friction dampers are realized slotting the upper part of the column above the 236 splice, adding cover plates and friction pads pre-s tressed with high strength pre-loadable 237 bolts on both web and flanges. To allow the gap ope ning, the slotted holes are designed to 238 accommodate a minimum rotation of 40 mrad [60], whi ch is the benchmark rotation 239 established by AISC 341-16 for Special Moment Frame s (SMFs). Similar provisions are given 240 in EC8 [3], which requires for Ductility Class High frames a rotation of 35 mrad. Between the 241 steel plates and the column, friction pads are inse rted. It is worth noting that the layout 242 presented in this paper is not explicitly consideri ng the possibility to accommodate a similar 243 rotation also in the weak direction, which is, inst ead, a situation rather common in practice. 244 Nevertheless, to provide a biaxial rotation capacit y to the connection it would be sufficient to 245 -9- oversize slightly the flange slots following the sa me simple geometrical rules used to over-size 246 the web holes. 247 a) Friction connections b) Re-centering system Fig. 2. Concept of the proposed solution 248 To provide a self-centering capability, pre-loaded threaded bars are introduced (Fig. 2b). 249 Additionally, to provide a sufficient deformability to the bar, a system of disk springs arranged 250 in series and parallel is installed in the assembly . 251 To assess the overall response of the connection (s ub-assembly of Fig. 3a), the behavior of the 252 whole system (connection, flange and web friction p ads, re-centering bars and column) can be 253 idealized by means of the simplified mechanical mod el delivered in Fig. 3b. The rotational 254 spring Cb accounts for the flexural stiffness of the cantile ver column of length equal to l0 (Fig. 255 3a), given by: 256 257  =3 (1) where Es is the steel modulus of elasticity, l0 is the column length up to the splice section and Ic 258 is the moment of inertia of the column profile. The translational spring Ff models the friction 259 pads on the column flanges. 260 -10- a) Joint subassembly b) Connection mechanical model c) Theoretical moment-rotation relationship Fig. 3. Concept of the proposed solution 261 The stiffness of this component can be assumed infi nite up to the achievement of the slip force 262 and equal to zero when this value is achieved. Simi larly, Fw models the friction pads on the 263 column web. The translational spring Ftb models the axial behaviour of the threaded bars 264 which work in series with the system of disk spring s, whose resistance is defined as F ds . The 265 stiffness of the threaded bars is given by: 266 267   =     (2) 268 and the stiffness of the disk springs is expressed as: 269 270  =    (3) 271 where nb is the number of bars employed in the connection s ymmetrically with respect to the 272 centroid of the column , n par is the number of disk springs in parallel, nser is the number of disk 273 springs in series and Kds1 is the stiffness of the single disk spring. Consid ering this mechanical 274 model, it is easy to verify that the typical moment -rotation behaviour of the connection can be 275 represented by a flag shape (Fig.3c). The moment M2 represents the decompression moment 276 which corresponds to attainment of the slippage for ce in all the friction pads. The first branch 277 of the moment-rotation curve is characterized by an infinite stiffness of the connection and, 278 therefore, the rotational stiffness of the whole sy stem is equal to KCb . The second branch, 279 -11- corresponds to the gap opening. In this phase, the slippage of the friction pads occurs, and the 280 rotational stiffness of the system is due to the st iffness of the threaded bars, disk springs and 281 column in bending, namely: 282 283 =     !  "# $ (4) 284 where hc is the column height. The branches 3 and 4 are cha racterized by the same stiffness of 285 the branches 1 and 2, respectively. The bending mom ent M0 represents the decompression 286 moment due to the sum of the axial load in the colu mn and to the pre-stress of the threaded 287 bars: 288 289 % =&'  +)*+  (5) 290 The bending moment M1 represents the contribution to the bending moment due to friction 291 pads, equal to: 292 293 %=',-ℎ − 0 1+'2+  (6) 294 where tfc is the thickness of the column flange. Considering these equations, it is easy to verify 295 that, from design point of view, the re-centering o f the connection can be guaranteed 296 imposing that: 297 298 % −%≥0 ⇒ '  ≥',-2− 0 + 1+'2−) (7) 299 3. DESIGN OF THE SPECIMENS FOR EXPERIMENTAL TESTS 300 301 With the set of equation previously reported, start ing from the definition of the design 302 actions, a column base connection has been designed . Owing to reasons of compatibility of the 303 specimen capacity with the available equipment, the axial load has been limited to the 25% of 304 the squash load, while the bending moment acting in the splice has been set equal to the 95% 305 of the plastic bending moment of the column. The sh ear load derives from the testing scheme 306 which is a cantilever representing, approximately, half column of the first storey of the 307 -12- building. Therefore, starting from a column profile HEB240, steel class S275, the following 308 design values have been calculated: 309 310 )= 8 ) 9 = 0,25 ) 9 = 728,75 >) 311 (8) 312 %= 198,5 >)A (9) 313 B=CD 9E= 128,1 >) 314 (10) 315 316 where l0=1,55 m is the distance between the force at the to p of the column and the splice 317 (Fig.3a), Npl is the column squash load, ν is the axial load ratio, Md is the assumed design 318 bending moment for the column base connection and Vd is the design value of the shear force. 319 Based on the shear design load Vd, firstly, the web component has been designed impo sing 320 that the slippage force on the web has to resist th e applied shear load. All plates are of S275 321 steel class. The friction pads have been chosen acc ording to the results of previous tests on 322 friction materials [62]. Basing on these results a friction coefficient μ=0,6, has been assumed. 323 Considering four bolts for both the upper and lower sides of the web connection, the pre-load 324 Fwp , for each bolt, has been determined as: 325 326 B= ' 2= F ' 2   ⇒ ' 2 = 26,7 >) 327 (11) 328 329 where Fw is the slip resistance of the web friction dampers , µ is the design value of the friction 330 coefficient, Fwp is the preloading force of the web bolts, nb is the number of web bolts and ns is 331 the number of friction interfaces (in this case, co nsidering the symmetrical configuration, this 332 is equal to two). Considering the design resistance , M14 HV bolts of 10.9 class have been 333 selected (HV stands for “Hochfeste Bolzen mit Vorsp annung”, which in English means “high 334 resistance bolts for pretension”). In order to desi gn the re-centering threaded bars, according 335 to Eq. (7), it has to be considered that the force in the bars depends on the slippage force of 336 the flange friction pads. Therefore, imposing the g lobal equilibrium between the internal and 337 external bending moment in correspondence of the sp lice, the following system can be 338 -13- written to design Ftb (the preloading force of the threaded bar) and Ff (the slip resistance of 339 the flange dampers) 340 341 H'  − ' ,-2 − 0 + 1 ≥ ' 2− )  '  + + ' , -ℎ − 0  1 = % −&'2+ ) * +  342 (12) 343 344 where hc is the column depth and tfc is the column flange thickness. For the sake of si mplicity, 345 if the lever arm of the friction force of the colum n flange friction dampers is approximated 346 with hc, the system of equations (12) leads to the followi ng simple design formulation: 347 348 '  ≥C" + − )  ⇒ '  ≥ 98,6 >) 349 (13) 350 Considering the design actions, for the specimens, two M20 threaded bars, having a maximum 351 capacity of 171,5 kN of pre-loading, have been adopted for the re-cente ring system. 352 Considering this capacity, the bar preload has been fixed equal to 100 kN. Therefore, system 353 (12), provides the following value of the design sl ippage force of the column flange friction 354 pads: 355 356 ',=C" + −  &'2+)+'  * ⇒ ',=298 ,9 >) (14) 357 Considering four bolts for both the upper and lower sides of the column flange connection, the 358 necessary pre-load Ffp , for each bolt, is: 359 360 ',=F ',   ⇒ ', =62 ,3 >) (15) 361 In this case, M20 HV bolts of 10.9 class have been selected. The last step of the design 362 procedure consists in the design of the disk spring s. Assuming a maximum rotation for the 363 joint equal to 40 mrad , the maximum gap-opening at the level of the re-ce ntering bar is 4,8 mm 364 (0.04x120 mm). Adopting standard disk springs with diameter equal to 45 mm, thickness 365 equal to 5 mm and height of the internal cone equal to 1.4 mm, three disk springs in parallel 366 -14- are necessary to resist to the bar yielding force. The resistance of each disk spring is about 80 367 kN, while the stiffness ( Kds1 ) is about 80 kN/mm. Considering the previously def ined 368 maximum displacement, Eq. (3) provides a minimum nu mber of 21 disk springs to be 369 arranged in sets of 3 springs in parallel (so-calle d “nested” configuration), 7 times in series 370 (so-called “back-to-back” configuration), leading t o an overall stiffness equal to Kds =35,36 371 kN/mm . 372 373 4. CYCLIC AND PSEUDO-DYNAMIC TESTS 374 375 The testing equipment is depicted in Fig. 4. Two ac tuators have been used: the first one, at the 376 top of the column is a MOOG Actuator (Maximum Load 3000 kN) governed under load control 377 in order to apply the axial load, the second one is an MTS 243.35 actuator, with a maximum 378 load capacity of 385 kN in compression and 240 kN i n tension and a piston stroke of 1016 379 mm, controlled under displacement control in order to apply a cyclic force at the top of the 380 column. 381 382 a) Side view b) Front view Fig. 4. Experimental layout. a) side view; b) fron t view 383 Regarding the measurement devices, a torque sensor Futek TAT430 has been used to 384 measure the initial torque applied to the bolts wit h the torque wrench, while four load cells 385 Futek LTH500 (capacity equal to 222kN) have been in stalled in the connection to monitor the 386 -15- tensile forces in the threaded bars and in two bolt s of one of the flange friction dampers (Fig. 387 5c). Additionally, LDT displacement transducers (ma x. 50mm) have been adopted in order to 388 measure the vertical displacements in both column s ides (Fig. 5c). Regarding the bolt 389 tightening procedure the initial pre-load, accordin g to EN 1090-2 specifications, was 390 increased of the 10% of the preload was added to th e bolt loads to account for random 391 variability of the bolt tightening and initial inst allation loss. Thus, a torque of 180 Nm was 392 applied in the flanges and 60 Nm in the webs, leadi ng to a force of 70kN and 30kN, 393 respectively. The pre-loading of the threaded bars was achieved by direct observation of the 394 load cells output. Concerning the actions, the axia l force was kept constant during the test, 395 while the cyclic horizontal load at the top of the column was applied consistently with the 396 loading protocol suggested by AISC 360-10. Four cyc lic tests have been performed varying the 397 axial load in the column, including or not includin g the re-centering bars. 398 399 (a) (b) (c) 400 Fig. 5. (a) Experimental layout; (b) Connection du ring the assembly; (c) view of the joint before the 401 test 402 In the different tests axial load ratios equal to 2 5% and 12,5% have been applied. The axial 403 loads were selected in a reasonable range of variat ion considering the typical size of MRFs 404 designed according to EC8. Specific values, in gene ral, obviously depend on the building plan 405 and frame configuration. Nevertheless, values rangi ng from 10% to 30% seem representative 406 of MRFs designed in DCH [60,63]. The adoption of a constant axial force is clearly not 407 reproducing the real loading situation of all the c olumns of a moment resisting frame. In fact, 408 due to overturning bending moments, especially the external columns of MRFs usually 409 undergo axial force variations during the earthquak e. The choice to adopt a constant axial 410 -16- force was done only to simplify the equipment used, as it is normally done in literature in 411 similar tests [64]. From the practical point of vie w, this situation better reproduces the 412 behavior of internal columns which, typically, unde rgo lower axial load fluctuations during 413 the seismic event. In the tests with lower values o f the column axial load, the total axial load in 414 the re-centering bars has been increased to 280 kN, which is still compatible with the pre- 415 loading capacity of the threaded bars but not suffi cient to respect Eq. (7) for guaranteeing the 416 flag shape behavior. In Table 1, a summary of the m ain values related to the loading condition 417 of the specimens is given. 418 Table 1. Main test data 419 Test Number Typology of Test Column Axial Load Total axial load in recentering bars Ratio between the applied load in recentering bars and the minimum one given by Eq.(7) Preloading of each web bolt Preloading of each flange bolt Residual rotation at the end of the test (Residual top displacement/l 0) kN kN kN kN [mrad] 1 Cyclic 728 (25% N p) 200 2.03 27 62 2.1 2 Cyclic 728 (25% N p) 0 - 27 62 4.1 3 Cyclic 365 (12.5% N p) 280 0.48 30 114 31.0 4 Cyclic 365 (12.5% N p) 0 - 30 114 49.7 5 Pseudo-dynamic 728 (25% N p) 200 2.03 27 62 1.7 6 Pseudo-dynamic 728 (25% N p) 0 - 27 62 5.2 7 Pseudo-dynamic 728 (25% N p) 200 2.03 27 62 2.7 In Figs. 6, the hysteretic curves of the experiment al tests are reported. In particular, in Fig. 6a 420 the experimental tests with higher axial load are d epicted, while Fig. 6b show the tests with 421 lower axial load ratio. The response of the connect ions reflected the expected behaviour, 422 highlighting in the different cases, the effect of the re-centering bar. In fact, from the results 423 presented in Figs.6, it can be clearly observed tha t the threaded bars have played an 424 important role. Tests 1 and 2 were carried out with the higher value of the axial load ratio 425 (25%), while test 3 and 4 were carried out with a r educed axial force (12.5%). In the first two 426 tests the self-centering behavior was expected (bec ause the size of the threaded bar was 427 defined considering an axial load ratio equal to th e 25%), while in the third and fourth test the 428 self-centering could not be achieved because the in itial tension in the bars was about half the 429 preload needed to achieve the theoretical self-cent ering condition. The 3 rd and 4 th tests were 430 carried out mainly to highlight the role of the re- centering threaded bar, even though in these 431 -17- cases to obtain a full self-centering, as already e videnced, higher capacity re-centering 432 systems should have been employed. The cyclic momen t-rotation curve of test 1 (Fig. 6a, red 433 line) highlights that the connection, with the axia l force considered in the design phase, was 434 able to return almost to the initial position with a very low value of the residual rotation (2.1 435 mrad), while in case of test 3 (Fig. 6b, red line) the residual rotation was higher (31 mrad) and 436 well beyond the constructional drift normally accep ted in the execution of steel structures 437 (usually lower than 5 mrad, depending on the number of columns and height of the building) 438 or the tolerance limit to be accepted accounting fo r the issues related to the building 439 functionality (which can be assumed accounting for the existing literature as equal to 5 mrad 440 as suggested in [65]). In any case, comparing Fig. 6a with Fig. 6b, the role of the re-centering 441 bars can be clearly noticed in both cases. 442 443 (a) Column Axial Load = 728 kN (b) Col umn Axial Load = 365 kN 444 Fig. 6. Moment-Rotation hysteretic curves of tested specime ns 445 446 Aiming to verify the ability of the proposed column base connection to dissipate energy and to 447 re-center, pseudo-dynamic tests (PsD) have been per formed at the Laboratory of Materials 448 and Structures of the University of Salerno. This t esting method combines an on-line 449 computer simulation of the dynamic problem (account ing for damping and inertial effects) 450 with experimental data regarding restoring forces a nd corresponding displacements due to 451 quasi-static application of loads, to provide reali stic dynamic response histories even in case 452 of non-linear behavior of severely damaged structur es [66]. Its main advantage is that it 453 adopts essentially the same equipment of a conventi onal quasi-static test, in which prescribed 454 load or displacement histories are imposed on the s pecimen by means of servo-hydraulic 455 actuators (Fig. 4). The structure to test has been idealized as a discrete-parameter system 456 consisting of one degree of freedom, controlled by the actuator. 457 -18- The classical equation of motion is solved by means of a direct step-by-step integration 458 scheme in which the mass and the viscous damping pr operties of the structures are modelled 459 analytically while the displacements, and consequen tly the restoring forces developed by the 460 structure, are measured with the external transduce rs positioned on a reference frame. 461 In the experimental test the MTS hydraulic actuator was used to apply the displacement 462 history to a system with a fictitious mass equal to 74t. The test was carried out neglecting any 463 additional viscous damping and applying a loading v elocity equal to 0.1 mm/s. 464 In order to perform the tests, the Kobe (Japan, 199 5) (record of the 16.1.1995, N-S direction) 465 and Spitak (Armenia, 1988) (record of the 12.7.1988 , N-S direction)earthquake records were 466 selected as ground motions. Record scale factors eq ual to 1.4 (PGA=0,35 g) for the Kobe 467 earthquake and equal to 1 (PGA=0.199 g) for the Spitak earthquake, were considered. The 468 selection of few earthquake records for a limited n umber of pseudo-dynamic tests is always, 469 under many point of views, arbitrary and cannot be representative of all the possible real 470 cases. In this activity, these two specific records were selected to compare earthquakes with 471 different features. In fact, as it can be noticed a lso from the response of the specimens, while 472 Kobe is a seismic event inducing a high number of l arge amplitude cycles, Spitak is 473 characterized mainly by two large reversal and many low amplitude cycles. The scale factor of 474 the seismic events was selected in order to achieve in the connection, approximately, a 475 rotation of 40 mrad. 476 In Fig.7 the moment-rotation plots of the pseudo-dy namic tests are reported. These pictures 477 confirm the improved performance of the proposed co lumn base connections in terms of 478 reduction of the residual rotations (Table 1). Also in this case, the comparison between the 479 moment-rotation curve of the column base connection with and without the re-centering 480 threaded bars (Fig. 8a) evidences the improvement o btained with the adoption of re- 481 centering bars. 482 -19- 483 (a) Kobe Earthquake (b) Spitak Earthquake 484 Fig. 7. Moment-Rotation curves of specimens submitted to ps eudodynamic tests 485 486 (a) Kobe Earthquake (b) Spitak Earthquake 487 Fig. 8. Displacement Time-history of specimens subm itted to pseudodynamic tests 488 This effect is also evidenced by the reduction of t he residual displacement after the simulated 489 earthquake (Table 1). In Fig.8 the time-history of the displacements at the top of the column 490 are shown for the three pseudo-dynamic tests. It ca n be observed that the column with the 491 proposed base connection with re-centering bars is characterized by residual displacement 492 after the earthquake always lower than 5 mrad [65]. 493 494 5. SIMULATIONS OF MRFs 495 496 In order to assess the effect of the adoption of th e proposed re-centering column base 497 connections over a structure, a preliminary time-hi story analysis of a MRF has been carried 498 out. The case study structure regards a four bays–s ix storeys scheme designed according to 499 the Theory of plastic mechanism control [67]. This methodology allows to select the column 500 size applying the upper bound theorem and the conce pt of mechanism equilibrium curve 501 ensuring the development of a failure mechanism of global type. The achievement of a failure 502 mechanism of global type appears very important in view of the application of a damage 503 -20- avoidance strategy adopting friction beam-column an d self-centering column base 504 connections, as suggested. 505 a) b) Fig. 9. a) Case-study frame. b) Scheme and Seismostruct model 506 507 The considered layout has inter-storey heights equa l to 3200 mm except for the first level 508 whose height is equal to 3500 mm, while the bays have all a span of 6000 mm. Regarding the 509 loads, a uniform dead load IJ= 4 >) A⁄ and a uniform live load MJ= 2 >) A⁄ (value given 510 by the code for residential buildings) have been co nsidered. Since the analyzed frame is the 511 perimeter frame of the building and the assumed tra nsversal bay span Lt is equal to 6000 mm, 512 a uniform dead load NJ= I J∙P! = 12.00 >) A ⁄ and a uniform live load RJ= M J∙P! = 513 6 >) A ⁄ have been considered, so that the design gravity l oad distribution has been 514 determined, in accordance with EC8, i.e. MS= 1.35N J+ 1.5R J= 25.20 >) A ⁄. With reference 515 to the seismic combination the load is determined a s MT= N J+ U RJ+   (where U is the 516 coefficient for the quasi-permanent value of the va riable actions, equal to 0.3 for residential 517 buildings) and, as a consequence, the applied reduc ed gravity load is MT= 12 + 0.3 ∙ 6 = 518 13.8 >) A ⁄. 519 6.00 m ANALYSED FRAME 6.00 m IPE 270 IPE 270 IPE 270 IPE 270 6.00 m 6.00 m 6.00 m HE 220 B qd = 25.20 kN/m 6.00 m 6.00 m 6.00 m 6.00 m 3.50 m 3.20 m 3.20 m 3.20 m 3.20 m 3.20 m IPE 270 HE 260 B HE 280 B HE 280 B HE 280 B HE 300 B IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 IPE 270 -21- a) b) Fig. 10. a) Beam-to-column connection with friction dampers; b) hysteretic response and 520 calibrated model 521 522 In the dynamic analyses, the seismic masses have be en evaluated starting from the uniform 523 distributed loads. The beam has been preliminary si zed considering the gravity load 524 combination (and afterwards checked also under seis mic loading conditions), leading to the 525 adoption of IPE270 profiles made of S275 steel. The column sections have been selected in 526 order to ensure a failure mechanism of global type, varying the shape at every storey from a 527 maximum of HEB 300 to a minimum of HEB 220 at the t op level of the building. The design 528 has been carried out considering a high seismicity region (PGA=0.35 g), soil type C, a seismic 529 response factor equal to 2.5, and a behavior factor equal to 6. 530 a) Mechanical model of the column base joint b) F-d laws of the adopted springs Fig. 11. Column base joint modelling implemented in Seismostruct [68] 531 -250 -200 -150 -100 -50 050 100 150 200 250 -0.08 -0.05 -0.03 0.00 0.03 0.05 0.08 Moment [kNm] Joint Rotation [rad] Hysteretic Curve Experimental Curve Calibration -22- 532 To assess the influence of the proposed connection over the global response, a first 533 comparison has been performed, modelling the frame with the software for dynamic analysis 534 Seismostruct [68] and analyzing the same MRF two ti mes: once assuming a fixed base and, 535 another time, introducing a set of springs able to accurately reproduce the typical behavior of 536 the proposed self-centering connection. In both cas es, the assumed beam-to-column 537 connections are the bolted joints with friction dam pers already tested in the European 538 research project FREEDAM, whose response is describ ed in [28] (Fig. 9). The hysteretic 539 behavior of the two connection typologies (column b ase and beam-to-column joint) has been 540 modelled by means of zero-length hysteretic links w hose mathematical parameters have been 541 properly calibrated by fitting the experimental mom ent-rotation response. In case of the 542 beam-to-column joint, a rotational spring has been used to model the connection, defining the 543 parameters by means of the optimization procedure d eveloped in [69]. In this case, to model 544 the hysteresis, the phenomenological model of Sivas elvan and co-authors [70] (commonly 545 called in the software Seismostruct “smooth”) has b een adopted. 546 547 a) N=600 kN b) N=500 kN Fig. 12. Comparisons between experiments and numeri cal model for different values of the 548 axial load 549 550 Conversely, to model the re-centering column base c onnections a complete mechanical model 551 based on the assembly of a set of springs able to r eproduce the response of the tests 552 previously reported, has been calibrated and verifi ed. This model has been calibrated in order 553 to grasp the effect of interaction between axial lo ad and bending moment over the response of 554 -23- the connection. The model is composed by four bilin ear springs in parallel with gap elements 555 able to simulate the uni-lateral hysteretic respons e of the friction dampers, plus a central 556 bilinear spring used to model the initial pre-stres s of the re-centering bar and the hysteretic 557 response of the bar itself (Fig.11). The yielding f orce of the springs modelling the friction 558 dampers is determined through Eqs.11 and 15, while the initial stiffness and the slope of the 559 hardening branch has been calibrated on the experim ental data, finding the optimized values 560 of K0=45 kN/mm and r= 0.01. Conversely, the stiffness of the spring model ling the re-centering 561 bars has been calculated as previously reported, na mely Kbar+ds =44.28 kN/mm. The yielding 562 force and the post-elastic stiffness of the recente ring bar have been calibrated starting from 563 the experimental data, defining Fy=250 kN and r= 0.01. With the selected parameters, the 564 application of the mechanical model to some test re sults is delivered in Fig.12. 565 In order to include in the structural model of the MRF the hysteretic behavior of the re- 566 centering connections, the parameters calibrated on the experimental tests have been 567 extended to model the particular configuration empl oyed in the reference building which, in 568 the specific case, represents the connection of an HEB 300 column. In any case, while some 569 model parameters are numerically calibrated by curv e fitting, the slip resistance of the friction 570 dampers and the pre-stress of the threaded bars hav e been re-calculated according to the 571 design procedure previously described in order to a chieve the full self-centering capacity of 572 the column base connection. 573 574 Fig. 13. Behavior of the designed base-plate connec tions for the case-study frame (Q=Force, 575 δ=displacement), evaluated on a shear length equal t o 1.55 m 576 577 The results of the design are summarized in Table 2 and Fig.13, for the external and internal 578 columns of the reference MRF. These two typologies need different design values of the pre- 579 -24- loading applied to the bolts and to the re-centerin g bars, because they are initially loaded with 580 different values of the axial force (250 kN for the external column, 500 kN for the internal). 581 Table 2. Model parameters for the re-centering connection ( notation explained in Fig.11b) 582 583 584 With the calibrated parameters, a time-history anal ysis of the MRF has been performed, 585 considering the first accelerogram of a set of eigh t natural records selected to match the EC8 586 reference elastic pseudo-acceleration spectrum (Fig .14). The fundamental natural period of 587 the structure has been assessed through a modal ana lysis determining a value of 1.6 seconds. 588 The two structures (fixed bases and self-centering connections) have the same period as the 589 proposed connection, due to the high initial stiffn ess, is nominally rigid. A damping ratio of the 590 5% was considered. 591 592 Fig. 14. Time-history response – top storey displacement (re d: with re-centering, black: 593 without re-centering) 594 595 00.5 11.5 22.5 33.5 0 0.5 1 1.5 2 2.5 3Sa [g] T [sec] EC8 EQ_1 EQ_2 EQ_3 EQ_4 EQ_5 EQ_6 EQ_7 EQ_8 Average spectrum Basic data of the selected ground motions ID Earthquake name Date Station name Sation Country Magnitude Fault mechanism PGA [g] EQ_1 Alkion 24.02.1981 Xylokastro - O.T.E. Greece 6.6 Normal 1.066 EQ_2 Montenegro 24.05.1979 Bar -Skupstina Opstine Montenegro 6.2 Reverse 0.492 EQ_3 Izmit 13.09.1999 Usgs Golden Station Kor Turkey 5.8 Strike-Slip 0.375 EQ_4 Aigion 15.06.1995 Aigio-OTE Greece 6.5 Normal 0.313 EQ_5 Umbria- Marche 26.09.1997 Castelnuovo- Assisi Italy 6.0 Normal 0.369 EQ_6 Izmit 17.08.1999 Istanbul - Zeytinburnu Turkey 7.4 Strike-Slip 0.480 EQ_7 Olfus 29.05.2008 Ljosafoss - Hydroelectric Power Iceland 6.3 Strike-Slip 0.315 EQ_8 Olfus 29.05.2008 Selfoss -City Hall Iceland 6.3 Strike-Slip 0.309 -25- The results given in Fig.15 highlight the enhanced response showing that with the proposed 596 column base joints it is possible to obtain a signi ficant improvement in terms of residual 597 drifts. In fact, while with traditional full-streng th column base connections the residual sway 598 displacement at the top of the building is equal to 350 mm (corresponding to 18 mrad of 599 average inclination of the column), with the employ ment of the proposed connections, it 600 reduces of about the 85%, achieving a residual top displacement at the end of the simulated 601 seismic event of 60 mm, corresponding to an average inclination of the building of about 3 602 mrad. This points out that, while with a traditiona l solution, the actual reparability of the 603 building would be compromised (18 mrad > 5 mrad), w ith the proposed self-centering 604 connections the residual drift reduces significantl y falling within the prescribed limits [65]. 605 606 Fig. 15 . Time-history response – top storey displacement (re d: with self-centering column 607 bases, black: without re-centering) 608 609 6. CONCLUSIONS 610 Aiming to obtain structures undergoing minimal dama ge able to return also to the original 611 configuration, in this paper, a new type of column base joint equipped with friction dampers 612 and re-centering threaded bars has been suggested. After presenting the conceptual design of 613 the connection, the results of tests and numerical simulations have been reported. The tests 614 carried out on the proposed connection under cyclic and pseudo-dynamic loading conditions 615 have evidenced that, according to the design assump tions, the threaded bars, if properly pre- 616 stressed, are able to work as elastic springs resto ring the connection to the initial 617 configuration. On the other hand, the numerical sim ulations of a case-study MRF have 618 evidenced that, compared to structures with classic al full-strength base plate joints, 619 structures with self-centering column bases can sel f-center returning to the initial 620 -26- configuration. These first analytical results seem promising and suggest extending the work to 621 other configurations in order to provide a wider va lidation of the system. 622 623 5. ACKNOWLEDGEMENTS 624 The research activity herein presented has been sup ported by the European Community by 625 research grant RFSR-CT-2015-00022. The support of t he European Commission within RFCS 626 Research & Innovation is gratefully acknowledged. 627