7 8 In this paper, the results of an experimental analysis regarding bea m-to-column joints 9 equipped with friction dampers is presented. Even though the overall concept is not new, the 10 connection structural detail and the friction pad material are different from previous 11 proposals. In particular, the beam is connected to the co lumn with a classical fixed T -stub 12 fastening the upper flange and a friction damper located at the beam lower flange. The friction 13 damper is composed of a stack of steel plates conceived to assure symmetrical friction. The 14 friction pads are made of steel plates coated with thermally sprayed aluminium . The friction 15 damper is designed in order to slide for a force level equal to or lower than the ratio between 16 the nominal flexural resistance of the connected beam and the lever arm, i.e. the distance 17 between the top T-stub and the friction damper. In this way, it is possible to obtain 18 connections able to diss ipate the seismic input energy almost without a ny damage to the steel 19 elemen ts, provided that all the joint components are designed with sufficient over -strength 20 with respect to the actions corresponding to the friction damper sliding force. In this paper, 21 such approach is validated reporting the results of an experimental campaig n. 22 23 Keywords: Friction d amper s, Connections, Free from yielding, Sprayed Aluminum 24 -2- 1. INTRODUCTION 25 26 The design of modern seismic resistant structures is based on a preliminary selection of the 27 zones which have to be designed to assure the dissipation of the earthquake input energy. 28 Dealing with Moment Resisting Frames (MRFs), the lo cation of such dissipative zones at the 29 beam ends is commonly preferred by adopting full-strength beam -to-column joints [1,2] 30 which have to be designed with suff icient over -strength with respect to the connected beams . 31 The required over -strength is aimed to assure the beam end yielding despite of the influence 32 of random material variability [3,4] and the amount of strain -hardening occurring before the 33 complete development of the ultimate flexural resistance of plastic hinges [5,6]. 34 To date, the classical design philosophy based on weak beam -strong column -strong joint 35 hierarchy has been widely appl ied in practical seismic design and surely provides some 36 advantages, such as the development of stable hysteresis loops [7-10] and the prevention o f 37 soft-storey mechanisms which have to be avoided because of their poor energy dissipation 38 capacity [11]. How ever, some drawbacks occur within the framework of the traditional design 39 approach. On one hand, the use of full -strength beam -to-column joints with the code required 40 over -strength can lead to the detailing of expensive structural connections which require the 41 use of continuity plates, supplementary column web plates, reinforcing ribs or cover plates or, 42 even, the use of haunched beams. On the other hand, also the overall frame design is costly, 43 because of the column over -strength required to fulfil the strength hierarchy criteria, 44 particularly in the case of long span beams , whose size is governed by gravity loads [12-15]. 45 In order to overcome the drawbacks related to the use of full -strength beam -to-column joints, 46 the use of partial -strength connect ions has been suggested and Eurocode 8 [1] has opened the 47 door to their use provided that their plastic rotation capacity is properly demonstrated. Such 48 design approach can be faced within the framework of the component approach [15-18]. The 49 fastening elem ents of the beam -to-column joints have to be properly designed by select ing the 50 weakest joint component , acting as dissipative component, and providing all the other joint 51 components with sufficient over -strength. Moreover, the weakest joint component has to be 52 designed to assure a ductile behaviour and the required plastic deformation capacity [19]. 53 In la st decades the application of partial strength joints to MRFs has been proposed and 54 supported by a high number of research programs, both theoretical and experimental, 55 devoted to characterise the behaviour of connections under monotonic [20-22] and cyclic 56 loading conditions [23-27]. Nevertheless, even though the effort provided by the scientific 57 community has been significant, there are still some issues which deserve further 58 -3- investigation, such as the codification of design criteria for dissipative joints or the 59 development of new types of dissipative connections easy to replace or not requiring 60 replacement after a severe seismic event [28-31]. 61 However, independently of the use of either full -strength or partial -strength beam -to-column 62 joints, the main draw back of the traditional seismic resistant design strategy is intrinsic in the 63 strategy itself. In fact, on one hand the structural damage is essential to dissipate the 64 earthquake input energy but, on the other hand, it is the main source of direct and indi rect 65 losses. For this reason, many researchers have focused their attention on the strategy of 66 supplementary energy dissipation with the aim to reduce the structural damage under 67 destructive seismic events and, as a consequence, the direct and indirect los ses. This strategy 68 is based on the use of energy dissipation devices which have to be inserted between couple s 69 of points of the structural scheme where high relative displacements or velocities occur under 70 the action of severe ground motions [33-36]. Such displacements or velocities are expected to 71 activate specifically designed passive energy dissipation systems based on simple 72 mechanisms such as hysteresis, friction or v iscosity of fluid s. 73 Starting from the background briefly summarized above, in order to overcome the drawback 74 of the traditional design approaches , research efforts have been recently devoted to the 75 practical devel opment of a new design strategy whose goal is the design of connections able 76 to withstand almost without any damage not only frequent and occasional seismic events, but 77 also destructive earthquakes such as those corresponding to rare and very rare events. 78 The concept behind this research is inspired to the strategy of suppleme ntary energy 79 dissipation, but it is based on the use of damping devices under a new perspective. In fact, 80 while passive control strategies have been commonly based on the integration of the energy 81 dissipation capacity of the primary structure by means of a supplementary dissipation coming 82 from damping devices; conversely, the new design strategy is based on the use of friction 83 dampers conceived in such a way to substitute the traditional dissipative zones of MRFs, i.e. 84 the beam ends. To this scope, beam -to-column connections can be equipped with friction 85 dampers which can be located either at the level of the two flanges [37-39] or at the bottom 86 flange level only [40-42]. Also the beam web to column flange connection can be equipped 87 with friction dampers. Moreover, symmetrical [39, 43 ] or asymmetrical friction devices can be 88 exploited [35, 38 ]. 89 -4- Figure 1: Comparison between different design strategies 90 In order to well clarify the aim of the work, its framework and the differences with either 91 traditional seismic design or supplementary energy dissipation strategy, the different 92 schemes are analysed in Fig. 1. In particular, Fig. 1a points out that dissipative zones of 93 traditional MRFs are located at the beam ends where plastic hinges have to be developed. The 94 seismic demand is usually expressed in terms of maximum interstorey drift ( 𝜗 in the figure) 95 which governs the plastic rotation expected in dissipative zones. T he supplementary energy 96 dissipation strategy (Fig. 1b) is aimed to the reduction of the seismic demand by introducing 97 seismic dampers which have to be located, for their effectiveness, between couple of points 98 subjected to high relative displacements. The supplementary energy dissipation provided by 99 such devices allows the reduction of the drift 𝜗 and, as a consequence, the reduction of the 100 structural damage occurring at the beam ends. Conversely, the substitution strategy (Fig. 1c) 101 allows the prevention o f the structural damage, because all the dissipative zones are 102 substituted by means of connections equipped with friction dampers. The expected drift 103 demand does not reduce when comparison is made with traditional structures (Fig. 1a), but 104 this drift lead s to very limited structural damage in some joints components , because the 105 rotation of the beam -to-column connections is accommodated by properly calibrating the 106 -5- stroke of the friction dampers (Fig. 1c). The maximum rotation allowed is practically given by 107 the ratio between the damper stroke and the lever arm, i.e. its distance from the centre of 108 rotation CR. However, it is useful to note that the relative displacement occurring between the 109 ends of the friction damper ( 𝛿=𝜗𝑑𝑏, being 𝑑𝑏 the beam depth) i s significantly less than the 110 one ( 𝛿=𝜗ℎ, being ℎ the interstorey height) occurring when the supplementary energy 111 dissipation strategy is applied (Fig. 1b). This is the main reason why cases a) and c) are 112 expected to provide similar drift demands. Finally , the best solution is obtained by combining 113 the substitution strategy with the supplementary energy dissipation strategy. Such combined 114 strategy (Fig. 1d) leads both to the reduction of the drift demands expected in case of 115 destructive seismic events and, in addition, to the prevention of significant damage in beam - 116 to-column connections . Obviously, the drift reduction is also an important benefit to re duce 117 damage to the building non -structural components. 118 Within the above framework , i.e. either substitution strategy or combined strategy , in this 119 paper a new beam -to-column connection equipped with friction dampers is investigated. In 120 particular, it is suggested to modify the classical detail of Double Split Tee Joints (DST) by 121 introducing a symmet rical friction damper at the level of the lower beam flange. With the 122 proposed connecting system , under bending actions, the joint is forced to rotate around the 123 upper T -stub, preventing slab damage, and the energy dissipation supply is provided by the 124 slippage of the lower beam flange on the friction pads which are made of steel plates coated 125 with thermally sprayed alumin ium. In this way, provided that the steel components of the 126 connection are properly over -strengthened, the joint resistance and the rotat ion capacity can 127 be easily governed by calibrating the preload applied to the frictional interfaces and realizing 128 slotted holes whose length provides an adequate stroke for the dissipative device. 129 As the overall concept of equipping beam -to-column connecti ons with friction dampers is not 130 new, it is important to underline the main differences with respect to similar connections 131 proposed by the New Zealand research group [39, 42, 44-45]. In particular, the main 132 differences regard the type of friction damper a dopted and the material used for the friction 133 pads . In fact, the connection structural detail herein presented is based on the use of 134 symmetric friction damper while the sliding hinge joint proposed by Clifton et al. is based on 135 an asymmetric friction damp er [39, 42]. The use of asymmetric friction dampers gives rise to 136 some yielding of the bolts which are subjected to bending moment, shear and axial force. Such 137 yielding is prevented when symmetric friction dampers are used like in the case of the beam - 138 to-column connections herein experimentally investigated. Moreover, the friction pads of the 139 connections herein investigated are made of steel plates coated by thermally sprayed 140 -6- aluminium; conversely, other re searchers have adopted brass or steel [38], bisalloy steel 141 grade [39, 42, 44 -45] or brake lining pads [37]. 142 In order to validate the proposed beam -to-column connection system , an experimental 143 program has been carried out at the Materials and Structures Laboratory of Salerno 144 University and the results obtained are herein presented. In particular, the results of 145 preliminary tests aimed at characteriz ing the frictional properties of new pads coated by 146 thermally sprayed aluminium are described and, afterwards, the results of two tests on re al- 147 scale external beam -to-column joints are presented. Furthermore, on the basis of the 148 component method , a design procedure -has been developed and adopted to design the 149 specimens The aim of the design procedure is to concentrate the energy dissipation i n the 150 friction damper and to prevent the yielding of all the other joint components . 151 152 2. EXPERIMENTAL ANALYSIS ON THERMALLY SPRAYED ALUMINUM INTERFACES 153 The results here presented constitute the further development of work already carried out in 154 a previous experimental project dealing with the characterisation of frictional interfaces made 155 of brass, steel and three different types of rubber [46, 47]. The results herein presented are 156 referred to a new type of friction pad composed by an 8 mm S275JR steel plate (steel 157 hardness 211 HBW, sand blasted surface ) coated with a thermally sprayed thin layer of 158 alumin ium (Fig. 2 ). The use of thermally sprayed aluminium coatings can lead to high value of 159 the friction coefficient [49], thus allowing the reduction of the size of the friction damper . 160 Even though not exhaustively, thermal spray can be defined as the application of coatings by 161 means of special devices/systems through which melted or molten spray material is propelled 162 at high speed onto a cleaned and prepared surface of the component to be coated. The coating 163 feedstock material is melted by a heat source. This liquid or molten material is then propelled 164 by process gases and sprayed onto a base material, where it solidifies and forms a solid layer. 165 There are several different processes to apply a thermal sprayed coating. The process applied 166 for the friction pads herein investigated is the electric arc wire spray. Even though the main 167 field o f application of sprayed alumin ium is the prevention of corrosion phenomena, the high 168 values of the friction coefficient obtained and the low cost of the raw material have already 169 suggested its use for friction pads in damping technology [48, 49]. 170 171 -7- a) Cross section of the coating b) Plan view of the surface Figure 2 : Cross -section of the coating and t exture of the tested m aterials 172 173 Figure 3 : Geometry of specimens and testing equipment 174 Sliding Plate -8- In order to evaluate the friction coefficient of interfaces made of steel plates sliding on 175 thermally sprayed aluminium such as those adopted in the design of the proposed joint, a sub - 176 assemblage composed by a stack of S275JR steel plates has been employed . The device has 177 been designed to allow the slippage of one of the inner steel plates on the interposed material. 178 To this scope, such plate is realised with slotted holes, while the other inner and the two outer 179 plates are realised with circular holes (Fig. 3). The clamping force is applied to the friction pad 180 by means of a maximum number of 8 high strength pre-loadable M20 bolts of 10.9 class [16, 181 50,51 ] and, in order to reduce preloading losses during the test, cone annular disc springs 182 have been used in substitution of the classical circular flat washers [46, 47 ]. The disc springs 183 adopted conform to standard DIN 6796. The tests have been carried out by means of a 184 universal testing machine Schenc k Hydropuls S56 (Fig. 3). In order to measure the axial 185 displacements the testing device is equipped with an LVDT, while the tension/compression 186 loads are measured by means of a load cell. The cyclic tests have been carried out under 187 displacement control f or different amplitudes at a frequency equal to 0.25 Hz. This value has 188 been selected to be compatible with the capabilities of the testing equipment. Even though 189 this frequency does not correspond to the values occurring under actual seismic events, it 190 seems from the available technical literature that the influence of the velocity is not 191 particularly significant. In fact, the sliding force versus velocity relationship can be modelled 192 by means of a power expression where the exponent is close to zero [60]. 193 The experimental analysis has been carried out in two phases. In the first phase , the influence 194 of the thickness of the coating layer on the value of the friction coefficient has been analysed 195 and the best value of the coating layer thickness has been se lected . To this scope, 196 experimental tests have been carried out on specimens with different values of the thickness 197 of the sprayed alumin ium coating . Successively, in the second phase, additional tests have 198 been carried out only on the selected interface to evaluate also the influence of the applied 199 pressure on the friction coefficient. 200 In particular, in the first phase, three different values of the coating thickness have been 201 considered (50 m, 150 m and 300 m) by adopting for each one , two different loading 202 sequences as reported in Table 1 . The torque has been assumed equal to 200 Nm and it has 203 been applied by mean s of a calibrated torque wrench, in order to achieve the desired bolt 204 preloading as described in Section 2.1. In the second phase, other two tests have been 205 performed only on the interface with coating thickness equal to 300 m considering other two 206 values of the tightening torque equal to 300 Nm and 400 Nm . The tightening was applied 207 consistently with the Europea n practice choosing among the possible available procedures 208 -9- (torque, combined, HRC or DTI) the torque method , which is the only one that allows to 209 tighten the bolts at values of the clamping force which are lower than the proof preload 210 corresponding to the 70% of the bolt nominal resistance. In fact, while with the combined 211 method, HRC bolts and DTI washers it i s possible to control only that the proof preload is 212 achieved, with the torque method intermediate value s of the clamping force can be applied 213 not losing accuracy. This is due, as explained in higher detail in Section 2.1, to the linear 214 relationship that e xists between torque and clamping force. More details on the accuracy of 215 the tightening procedure when applied to friction dampers are given in [61]. 216 Table 1: Summary of the tests carried out in the experimental programme 217 Coating Thickness Number of tightened bolts Bolt Torque Number of Cycles of the loading sequence Amplitude 50 m 4 200 Nm 10 +/-15 mm 8 45 150 m 4 200 Nm 10 +/- 15 mm 8 45 300 m 4 200 Nm 10 +/- 15 mm 8 200 Nm 20 4 300 Nm 10 4 400 Nm 10 2.1 TESTS ON THE COMPONENT 218 The main goals of the experimental campaign are, on one hand, the evaluation of the friction 219 coefficient for different values of the normal force acting on the sliding interface and, on the 220 other hand, the assessment of the cyclic response in terms of stability of the cycles obtained 221 and energy dissipation capacity. In the following the test results are discussed reporting the 222 values obtained for the friction coefficient , determined as: 223 𝜇=𝐹 𝑛𝑠𝑛𝑏 𝑁𝑏 (1) where ns is the number of surfaces in contact, nb is the number of bolts, 𝑁𝑏 is the bolt 224 preloading force and 𝐹 is the sliding force. In particular, the bolt preloading force 𝑁𝑏 is 225 determined starting from the knowledge of the tightening torque by means of the following 226 expression: 227 𝑁𝑏=𝑇𝑏 𝑘 𝑑𝑏 (2) where Tb is the value of the tightening torque , k is a factor accounting for the friction arising 228 between the bolt nut and the plate and between the threads of the bolt shank and nut and db is 229 the bolt nominal diameter. In agreement with [54] the value of k has been assumed equal to 230 0.20. 231 -10- Even though it is common practice to derive the friction coefficient of such devices by means 232 of Eq. (1), it is useful to note that, according to EN 1993 -1-8 [18], the design slip resistance has 233 to be computed as: 234 𝐹𝑠,𝑅𝑑=𝑘𝑠𝑛𝑠𝑛𝑏𝑁𝑏 𝛾𝑀3 (3) where 𝛾𝑀3 is the partial safety factor, to be assumed equal to 1.25 for ultimate limit state s and 235 equal to 1.10 for serviceability limit stat es, and ks is a coefficient accounting for the hole 236 typology. In particular, in case of long slotted holes (i.e. with clearance on the length at least 237 equal to 1.5d [55]) with the axes of the slot parallel to the direction of load transfer, the value 238 ks=0.63 is sugg ested. 239 240 2.1.1 Tests with variable thickness of the coating layer 241 In order to evaluate the slip resistance of alumin ium sprayed surface s, sliding on steel, 242 characterized by a different value of the coating thickness , different tests have been carried 243 out. As a synthesis of the results, some of the tests and the values of the friction coefficient 244 versus the cumulated displacement are delivered in Fig. 4 , with reference to the group of 245 specimens of the 2nd loading sequence . The results obtained are summarized in Table 2 . 246 1st Loading Sequence -0.6-0.4-0.200.20.40.6 0 100 200 300 400 500 600 700FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -4Bolts -Ts=200 Nm (50 m) -0.6-0.4-0.200.20.40.60.8 0 100 200 300 400 500 600 700FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -4Bolts -Ts=200 Nm (150 m) -11- 2nd Loading Sequence Figure 4 : Results obtained with coatings having different thickness 247 The results of the tests point out that the behavio ur of all the interfaces is quite stable . All the 248 interfaces exhibited a high initial stiffness and, after reaching the sliding force, began to slide 249 -0.6-0.4-0.200.20.40.60.8 0 100 200 300 400 500 600 700FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -4Bolts -Ts=200 Nm (300 m) -0.6-0.4-0.200.20.40.6 0 500 1000 1500 2000 2500 3000FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -8Bolts -Ts=200 Nm (50 m) -0.6-0.4-0.200.20.40.6 0 500 1000 1500 2000 2500 3000FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -8Bolts -Ts=200 Nm (150 m) -0.6-0.4-0.200.20.40.6 0 200 400 600 800 1000 1200 1400FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -8Bolts -Ts=200 Nm (300 m) -12- with a force similar to the initial one, evidencing that there is not a significant differ ence 250 between the static and kinetic coefficient of friction. 251 It useful to note that the displacements given in Fig. 4 are total displacements including also 252 the elastic contribution due to the axial stiffness of the plate elements constituting the 253 specime ns. 254 In addition, after the first cycle, all the subsequent cycles reached a value of the maximum 255 force slightly lower than the initial one, evidencing the role played by the wearing of the 256 friction material of the coated surface ( Fig. 5 ). However, regarding the degradation of the slip 257 resistance occurring as far as the cumulated displacement increases, it is useful to point out 258 that, in case of destructive seismic events, the friction dampers equipping beam -to-column 259 connections are expecte d to be subjected to cumulated displacement of about 250 mm [56], 260 so that the degradation pointed out in Fig. 4 can occur only in case of repeated earthquakes. 261 Analyzing the results summarized in Table 2 , it is possible to observe that, considering all the 262 tested interfaces, the value of the friction coefficient ranges in between 0.37 and 0.59, 263 revealing that thermally sprayed aluminium is able to provide a value of the coefficient of 264 friction higher than the metallic and rubber interfaces tested in past ex perimental 265 investigations [46]. In particular, the highest value of the friction coefficient was obtained 266 with the thickness of 150 m and the lowest with the thickness of 50 m. 267 In addition, the reported tests show that the degradation of the sliding force significantly 268 depends on the thickness of aluminium coating. In fact, considering the first loading sequence 269 (Table 2), in case of 50 m and 300 m thickness, the degradation of the sliding force is very 270 low and the ratio between the initial and final friction coefficient is equal to 1.12 and 1.02 271 respectively. In the first case, probably the result is due to the lower sliding force that 272 provided a limited consumption of the fric tion pad. Conversely, a higher degradation was 273 shown by the 150 m interface, whose initial/final sliding force ratio was equal to 1.22. 274 Therefore, analyzing the results on these tests, for the subsequent phase of the experimental 275 work, only the coating wi th thickness of 300 m has been considered. In fact, as a result of the 276 first phase tests, this thickness of the coating layer is believed to provide a good compromise 277 between the initial value of friction coefficient and degradation of the sliding force d uring a 278 cyclic loading history. 279 -13- 280 Figure 5 : Appearance of surface of f riction pads after test ing 281 282 Table 2: Friction coefficients obtained for different coating thickness 283 Thickness of the coating 1st loading sequence 2nd loading sequence Initial Final Initial Final 50 m 0.37 0.33 0.37 0.31 150 m 0.59 0.48 0.52 0.29 300 m 0.52 0.51 0.45 0.37 284 Regarding the values of the friction coefficient obtained, it is useful to note that the values 285 given in Table 2 could seem comparable with those which can be obtained by means of shot 286 blasted steel, so that the incentive in using sprayed aluminium could seem not much. As an 287 example, a recent experimental program dealing with the design of slip resistance lap joints 288 [57] has shown , in case of normal holes, a friction coefficient equal to 0.56 in case of joints 289 made of S275 steel whose surface is blasted with shot steel to Sa 21/2 and equal to 0.52 in case 290 of joints made of S275 steel whose surface is blasted with shot steel to Sa 3 and spray - 291 metalized with 75 m zinc. However, it has to be considered that the values given in Table 2 292 are obtained for long slotted holes, so that they are also the result of a smaller contact area 293 leading to higher local pressures. The influence of the hole typology is accounted for by means 294 of the coefficient ks in Eq. (3) which, according to Eurocode 3, is equal to 0.63 aiming to 295 provide a safe side value of the slip resistance. According to [58], such influence is about 30% 296 in case of long slotted holes, so that t he value 0.59 obtained in case of 150 m coating 297 thickness is equivalent to a value of about 0.77 for normal holes. This consideration points out 298 the advantage which can be obtained by mean s of thermally sprayed aluminium. 299 -14- Moreover, it is useful to note that values of the friction coefficient higher than those obtained 300 in the experimental tests herein presented have been obtained by other researchers [48]. This 301 difference can be attributed to the differences in the preliminary treatme nts applied before 302 thermal spraying and/or to t he main setting parameters of the machine employed to apply 303 the coatings , i.e. the amperage, voltage and pressure of the air, which have to be changed 304 according to the melting temperature of the materials appl ied. 305 306 2.1.2 Tests with variable pressure 307 308 As aforesaid, in order to evaluate the influence of the tightening pressure on the value of the 309 slip resistance, additional tests have been carried on further specimens having the selected 310 coating thickness. In particular, three different values of the tightening torque have been 311 considered and only one loading sequence of ten cycles with amplitude of +/ - 15 mm has been 312 adopted. The behaviour exhibited by such specimens is reported in Fig. 6. It evidences a 313 respo nse very similar to that reported in previous tests, but reveals that the value of the initial 314 coefficient of friction, for this interface, is also affected by the pressure applied on the 315 interface. In fact, as it is possible to observe, as far as the valu e of the tightening torque 316 increases, the initial friction coefficient decreases (Fig. 7). 317 318 (Ts=200Nm) (Ts=300Nm) (Ts=400Nm) -400-300-200-1000100200300400 -20 -15 -10 -5 0 5 10 15 20FORCE [kN] DISPLACEMENT [mm]TS Aluminum on Steel -4 Bolts -Ts=200 Nm (300 m) -400-300-200-1000100200300400 -20 -15 -10 -5 0 5 10 15 20FORCE [kN] DISPLACEMENT [mm]TS Aluminum on Steel -4 Bolts -Ts=300 Nm (300 m) -400-300-200-1000100200300400 -20 -15 -10 -5 0 5 10 15 20FORCE [kN] DISPLACEMENT [mm]TS Aluminum on Steel -4 Bolts -Ts=400 Nm (300 m) -15- Figure 6: Cyclic response for different values of the applied tightening torque 319 Figure 7: Friction coefficient vs. cumulated displacement for different tightening levels 320 321 Therefore, in order to develop a mechanical model to design beam -to-column connections 322 equipped with friction dampers with thermally sprayed alumin ium pads , a relationship 323 providing the initial friction coefficient as a function of the pressure applied at the interface 324 has been derived . To this scope, the pressure applied at the interface has been estimated 325 assuming that the normal force is transferred by th e bolts to the p lates spreading with a slope 326 equal to 45°, obtaining the following equation which includes the influence of the long slotted 327 hole ( Fig. 8 ): 328 329 𝑝=4𝑁𝑏 𝜋2𝑑𝑏2(1−2arcsin ()⁄ 𝜋− 𝜋√1−( )2 )2 (4) 330 where 𝑑𝑏 is the slot width and: 331 =𝑑𝑏ℎ+2(𝑡𝑤+𝑡𝑝+𝑡𝑓𝑝) 𝑑𝑏 (5) 332 -0.8-0.6-0.4-0.200.20.40.60.8 0 100 200 300 400 500 600FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -4 Bolts -Ts=200 Nm (300 m) -0.8-0.6-0.4-0.200.20.40.60.8 0 100 200 300 400 500 600FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -4 Bolts -Ts=300 Nm (300 m) -0.8-0.6-0.4-0.200.20.40.60.8 0 100 200 300 400 500 600FRICTION COEFFICIENT CUMULATED DISPLACEMENT [mm]TS Aluminum on Steel -4 Bolts -Ts=400 Nm (300 m) -16- being db the bolt diameter, dbh the bolt head diameter, tw the washer thickness, tp the outer 333 plate thickness and tfp the friction pad thickness . 334 335 Figure 8: Assumed geometry of the contact pressure area 336 It is useful to note that the above relationships are valid provided that the distance between 337 the slots and the bolt pitch is greater than db and, in addition, the distance between the slot 338 and the plate edge is greater than db /2. When such conditions are not satisfied the 339 interaction has to be accounted for. In Fig. 9 the result of the regression analysis of the 340 experimental data is reported . Nevertheless, being referred only to three experimental data, 341 this regression has to be considered with some caution. 342 The influence of the pressure on the friction coefficient is justified considering that all 343 surfaces are rough on a microscopic scale an d real contact is obtained over a small fraction of 344 the apparent contact area. Friction is related to the real contact area and independent of the 345 apparent contact area. Real contact area is affected by the applied pressure which, therefore, 346 affect s the re sulting friction coefficient. In particular, the increase of the applied pressure 347 tends, on one hand, to flatten the surface asperities, thus reducing the amount of friction due 348 to the mechanism of asperity interlocking and, on the other hand, to increase the real contact 349 area thus improving the adhesion mechanism. T he prevailing between these two phenomena 350 will govern the dependence of the friction coefficient on pressure and can be investigated by 351 experimental results only. 352 External PlateFriction PadExternal Plate Sliding PlateFriction Pad 45°45°tptw tfp Long Slotted Holedb dbContact Pressure Area -17- 353 Figure 9: Influen ce of the contact pressure on the friction coefficient 354 355 3 CONCEPT AND DESIGN OF TESTED SPECIMENS 356 Starting from the above results , the design of dissipative connections equipped with friction 357 dampers was performed. The joint detail proposed in this paper is the modification of the 358 classical detail of a double split t ee beam -to-column connection obtained by connecting the 359 beam bottom flange to the column flange by means of three steel angles, composed by 360 welding , and by interposing between them and the beam flange one or two friction pads (Fig. 361 10). In this way , the energy dissipation supply is provided by the friction damper, while all the 362 other element s of the connection and the beam are designed in order to be free from yielding . 363 Exception has to be made for the stem of T-stub and angles , close to the T -stub flange and 364 vertical angle leg respectively, which in order to accommodate the rotation are subjected to 365 minor yielding. This st rategy , as already demonstrated in some prelim inary studies on similar 366 prototypes, allows the development of beam -to-column connections with high energy 367 dissipation able to accommodate the required rotations without any significant damage [47]. 368 The classical T -stub connection at the top flange level is aimed to fix the cen tre of rotatio n 369 with the goal of preventing the damage to the reinforced concrete slab, while the bottom 370 damper provides, thanks to its stroke, the required rotation capacity and the energy 371 dissipation. 372 A significant advantage of this joint configuration is that, by con trolling the tightening torque 373 applied to the bolts, it is possible to calibrate magnitude of the sliding resistance of the friction 374 damper and, therefore, depending on the lever arm, the magnitude of the bending moment 375 y = -0.0116 p + 0.6537 R² = 0.9989 [p (MPa)] 0.000.100.200.300.400.500.60 0 5 10 15 20 25 30 35 40FRICTION COEFFICIENT Pressure [MPa]TS Aluminum on Steel Interface -Dynamic Friction Coefficient -18- transmitted to the column. It is use ful to note that, in order to assure that the beam end 376 remains in elastic range, the bending sliding resistance of the connection has to be not greater 377 than the nominal bending resistance of the connected beam. However, in order to maximize 378 the exploitatio n of the beam under gravity loads and under the internal actions occurring 379 when the structure is subjected to horizontal forces corresponding to serviceability limit 380 states , such as wind actions or frequent seismic events having a low return period, the ra tio 381 between the bending sliding resistance of the connection and the nominal bending resistance 382 of the connected beam has to be as close as possible to 1.0. In order to satisfy the above design 383 criterion and to maximize the exploitation of the connected be am, the use of a beam end 384 haunch to increase and calibrate the lever arm can also be suggested (Fig. 11). In such a way, 385 the beam section is fully exploited, but both the oversizing of the other joint components 386 (such as the panel zone usually requiring su pplementary web plates, the beam end requiring 387 reinforcing ribs or cover plates , the increase of the bolt diameter, etc.) and the column 388 oversizing (because of beam -column hierarchy criterion) can be significantly limited and/or 389 controlled. 390 391 Figure 10: Beam -to-column connection equipped with friction dampers 392 393 -19- 394 Figure 11: Beam -to-column connection equipped with friction dampers an d additional haunch 395 396 Within the experimental activity described in this paper two different joints equipped with 397 the friction damper previously tested un der uniaxial loading conditions were tested under 398 cyclic loading conditions . In particular, the two joints fasten an IPE270 beam to an HEB 200 399 column , both of them made of S275 steel . They are identified as follows: 400 • TSJ-SA300 -320 -CYC 12: it is a joint where the upper beam flange is connected to the 401 column b y means of a T -stub, while the lower beam flange is fastened to the column by 402 means of a system of three angles (Fig. 10). Three steel plate s with a 300 m coating of 403 thermally sprayed alumin ium are located in between the beam flange and the angles in 404 order to realize the friction damper; 405 • TSJ-H-SA300 -260 -CYC 13: it is a joint with a detail v ery similar to the previous one where, 406 in order to increase the lever arm, the friction damper is applied by means of an additiona l 407 haunch which is welded to the beam. This allows the reduction of the tightening torque 408 and/or of the number of bolts (Fig. 11). 409 Starting from the values of the friction coefficients derived in previous section, the design of 410 the dissipative connections equipped with friction pads was performed. Within the 411 framework of the component method, the following components have to be designed: the 412 shear panel, the column web panels in tension and compression, the T -stub/ angles and the 413 friction damper. In order to obtain a joint where the only component providing energy 414 dissipation is the friction damper, the steel parts have to possess sufficient overstrength w ith 415 respect to the maximum force that the friction damper is ab le to transmit. According to this 416 hierarchy, the geometry of all the elements composing the joint was defined by exploiting the 417 formulations provided by literature models or by means o f the formulations given in 418 Eurocode 3 [17]. 419 -20- Specimen TSJ -SA300 -320 -CYC 12 was designed to develop a bending moment evaluated at the 420 level of the column face equal to the beam nominal plastic resistance (133 kNm). It means that 421 the joint was designed to assure that the beam end remains in elastic range. In fact, 422 considering that the friction damper is characterized by a rigid -plastic response , imposing 423 that the slip resistance of the damper occurs at the attainment of a bending moment equal t o 424 the characteristic flexural resistance of the beam , assures that the connected member remains 425 elastic . A similar criterion has been adopted for specimen TSJ -H-SA300 -260 -CYC 13 426 accounting , however, for the possibility to increase the slip resistance of the connection 427 beyond the nominal resistance of the beam evaluated at the column face. In fact, in this case, it 428 is possible to account for the increase of resistance of the beam in the haunched part 429 increasing , consequently, the slip resista nce of the damper . Therefore , this joint has been 430 designed to develop a bending moment equal to 1.4 times the beam nominal resistance , 431 namely 187 kNm . Taking into account that the beam length (i.e. the distance between the 432 actuator and the column face) is equal to 1460 mm and that the un-strengthened segment of 433 the beam is equal to 885 mm, when the design bending moment at the column face is achieved 434 (187 kNm), the bending moment acting at the end of the haunch is equal to 435 187x885/ 1460 ≅113 kNm (which is about 0.85 times the beam nominal resistance) . It means 436 that also for this specimen, when slippage occurs, the bending moment in the un -strengthened 437 portion of the beam is lower than the beam nominal resistance , assuring that it remains 438 elastic . These reference values of the flexural resistance were used to define the tightening 439 torque of the bolts of the friction damper accounting for the friction coefficient values 440 previously reported. 441 All the remaining joint components have been designed with an appropriate overstrength. In 442 particular, the column shear panel has been reinforced with a couple of 10 mm 443 supplementary plat es welded on the column web, while the panels in tension and 444 compression have been stiffened with a couple of continuity plates with a thickness equal to 445 the beam flange thickness. However, in order to obtain structural details which are more cost 446 effectiv e the possibility to omit these strengthening elements can be evaluated. In addition, 447 the upper T -stub and the lower angles have been designed in order to remain in elastic range 448 under a force level corresponding to the friction damper design force. 449 Deali ng with the friction damper, the parameters to be controlled in the design are the length 450 of the slots made on the beam flange or on the haunch (which allow the slip of the beam flange 451 on the friction pad) and the tightening torque to be applied to the bol ts which are used to 452 preload the friction pads. In particular, the first parameter governs the rotational capacity of 453 -21- the connection and, as a consequence, it can be selected starting from the expected value of 454 the joint rotation ( ) under the most severe design earthquake, by using the following 455 equation: 456 () z dp n Lb br slot 2 1++−= (6) where db is the bolts diameter, nbr is the number of bolt rows used to fasten the stem of the 457 angles to the beam bottom flange, p is the bolt pitch and z is the lever arm. For the tested 458 specimens db=20 mm, n br=4, p=60 mm and z = 265 mm for specimen TSJ-SA300 -320 -CYC 12 459 and z = 411 mm for specimen TSJ-H-SA300 -260 -CYC 13, respectively. Therefore, by assuming 460  =0.08 rad , according to this criterion a length of the slot equal to 270 mm has been chosen 461 for the two joints. The clear gap between the beam edge and the inside o f the flange of the Tee 462 section has been assumed equal to about two times the thickness of T -stub stem to allow the 463 local yielding required to accommodate the beam rotation. 464 The second parameter governs the magnitude of the bending moment that the joint is able to 465 withstand before the slippage of the friction damper occurs. In particular, the design force o f 466 the friction damper can be easily defined as the design bending moment divided by the lever 467 arm. The design friction resistance of the damper device is equal to 133/0.265 ≅502 kN in 468 case of specimen TSJ -SA300 -320 -CYC12 and equal to 187/0.411 ≅455 kN in cas e of specimen 469 TSJ-H-SA300 -260 -CYC 13. S tarting from the knowledge of the friction damper design force, 470 exploiting Equations ( 1) and (2) and the relationship between the friction coefficient and the 471 pressure applied at the interface (Fig. 9), it is possible to determine the value of the tightening 472 torque to be applied , equal to 320 Nm and 260 Nm for the two tested specimens. The details 473 of the specimens are deliver ed in Fig. 12. 474 475 476 Figure 12: Details of the tested specimens. Left: TSJ -SA300 -320-CYC12; Right: TSJ -SA300 -260-CYC13 477 478 4560266045105 105 45 17645358045 86 4517645 114 4520045 70 70 70 60 60 60 60 60 350350 75IPE 270HEB 200 86 45Supplementary web plates t = 10 mm Friction pads t = 8mm Continuity plates t = 10 mmSlotted holesAngles - t w = 15mmT-stub - t w = 20mm4 Bolts M27 8.8 Class 8 Bolts M20 8.8 Class 75 60 60 60 6045 70 70 70 60 35035045 86 45176HEB 200 IPE 27045358045200 111 454560 6045236 17686 45 45Haunch 146411 Slotted holesFriction pads t = 8mmAngles - t w = 15mmT-stub - t w = 20mm4 Bolts M27 8.8 Class 8 Bolts M20 8.8 Classt=25 mm t=25 mm hb=270mm hh=140mm -22- 479 4 EXPERIMENTAL LAYOUT 480 The e xperimental tests were carried out at Material s and Structures Laboratory of Salerno 481 University. Two steel hinges, designed to resist shear actions up to 2000 kN and bolted to the 482 strong floor base beam, were adopted to connect the specimens to the reacting system. The 483 specimen s were assembled with the column (HEB 200) in horizontal position, connecte d to 484 the hinges, and the beam (IPE 270) in vertical position ( Fig. 13 ). The loads were applied by 485 means of two different hydraulic actuators. The first one used to apply, under force control, an 486 axial load in the column equal to 30% of the squash load . The second actuator to apply, under 487 displacement control, the desired displacement history at the beam end. In order to avoid the 488 lateral -torsional buckling of the beam , an horizontal transversal frame was conceived to work 489 as a guide which restraints th e lateral displacement of the beam . The loading history was 490 defined in terms of drift angle, according to the testing protocol provided by AISC, 2005 [59]. 491 492 Figure 13: Experimental layout for tests on joints 493 During the tests the displacement history imposed by the top actuator and the displacements 494 of the different joint components were acquired In particular, in each test, four LVDT were 495 used to measure the local displacement of the friction damper, the rotation of the panel zone 496 and to control that the axial displacements exhibited by the fixed T -stub were negligible ( Fig. 497 14). 498 Aiming at the evaluation of the beam end displacements due to the beam -to-column joint 499 rotation onl y, the top displacements of the cantil ever beam measured by means of the LVDT 500 equipping the actuator were corrected by subtracting the elastic contribution due to the beam 501 and column flexural deformability according to the following relationship: 502 -23-     ++    +−−=a La a LL EILFL EIFL c cc cbc bb MTS j26 2 12 322 3  (7) 503 504 Figure 14: Position of LVDT s used to measure the local displacements 505 506 where MTS is the overall top horizontal displacement evaluated at the level of MTS actuator, Ib 507 and Ic are the beam and column inertia moments, Lc is the column length, Lb is the beam length 508 and a is the length of the rigid parts, due to the steel hinges (Fig. 15 ). 509 a) beam contribution b) column contribution c) joint contribution 510 Figure 15: Scheme for e valuating the joint contribution s to the overall top displacement 511 512 The identification label of the tests refers to the following issues : 1- Joint typology, i.e. Tee 513 Stub Joint (TSJ) or Tee Stub Joint with haunch (TSJ -H) / 2 – Friction interface, i.e. SA300 514 (Sprayed Alumin ium with 300 m coating ) / 3 – Tightening torque of the bolts in Nm / 4 – 515 Progressive number of cyclic tests, i.e. CYCXX. 516 517 t c n t c n t c na=350 mm a=350 mm Lb=1560 mm Lc=2000 mm -24- 5 CYCLIC BEHAVIOR OF SPECIMENS 518 As aforesaid , the two joints (Fig. 16) were tested under cyclic loading conditions follow ing the 519 drift history suggested by the AISC loading protocol. As expected , on the basis of the design 520 criteria adopted, in all the experimental tests there has not been any significant damage of the 521 joint compon ents, but only the wearing of the friction pads. Therefore, a very important 522 outcome of the experimental program is the verification that this connection typology can be 523 subjected to repeated cyclic rotation histories, i.e. to repeated earthquakes, by only 524 substituting the friction pads, if needed, and by tightening again the bolts to reach the desired 525 preloading level. In addition, the rotation capacity can be easily calibrated by simply 526 governing the length of the slotted holes. In fact, in all the tests the rotation demand applied to 527 the joint has been completely due to the slippage of the friction damper located at the bottom 528 beam flange . The experimental results are in line with the outcomes of the tests on the friction 529 interfaces pointing out that, as expected, the cyclic behaviour of the joint is mainly governed 530 by the cyclic behaviour of the weakest joint component, i.e. the friction damper. 531 532 Figure 16 : Specimens before testing 533 In fact, in both tests the response has been very similar to that exhibited during the uniaxial 534 tests on the interface. At low force values the joint s exhibited an elastic behaviour 535 characterized by a high initial stiffness . In particular, in case of joint TS-H-SA300 -260 -CYC13 536 the initial stiffness was sensibly higher due to the increased lever arm provided by the 537 additional haunch . When the force applied at the end of the cantilever beam reached a value 538 approximately equal to the design bending moment divid ed by the beam length , the slippag e 539 of the friction dampers started (Figs. 17 -18). After the first slippage, the joint s hysteretic 540 response was characterized for all the loading/unloading cycles approximately by a 541 -25- parallelogram shape with a slight strength degradation as the number of the cycles of the 542 loading history increases (Fig. 19 ). 543 By the comparison with the results of the tests carried out on the friction dampers (i.e. the 544 component alone) , it is worth noting that the main difference between the hysteresis cycle s of 545 the friction DST joint and those observed during the uni axial tension test is the asymmetric 546 response . This difference is mainly due to the role played by the beam rotation in the 547 kinematic mechanism. In fact, the beam rotation causes two effects that give rise to an 548 increase of the bending moment as far as the beam rotation increases. On one hand, there is 549 an increase of the local pressure on the friction pads due to the reaction force provided by the 550 stem of the fixed T -stub at the top beam flange level and by the stems of the angles at the 551 bottom flange level, that behave in a way similar to a pocket foundation. Because of this 552 pocket foundation effect, the increase of the pressure on the friction pads depends on the 553 direction of the horizontal displacement of the cantilever. Therefore, the asymmetry of the 554 cyclic response is due to the reduction of the friction coefficient when the pressure on the 555 friction pads increases. Additionally, when the dampers are in tension a deformation of the 556 steel angles occurs , probably providing a reduction of the bolt forces due to the opening of a 557 small gap in the web plates. On the other hand, minor yielding of the tee stems at the stem -to- 558 flange connection of top T -stub and bottom angles contributes to the total bending resistance 559 providing a slight hardening behaviour as experimentally observed (Fig. 19). 560 561 Figure 1 7: Sliding motion of the friction damper 562 -26- 563 Figure 18 : Force -Displacement response of the friction dampers 564 The most important feature of the proposed connection is that , as confirmed by the 565 experimental results , it is able to provide a high dissipative capacity also under values of the 566 rotation significantly greater than the minimum value , equal to 35 mrad , required by 567 Eurocode 8 [1] for frames in High Ductility Class. Furthermore, it is possible to observe from 568 Fig. 18 a nd Fig. 19 that the resistance s of the joints and of the dampers approximately 569 correspond to the design one s, confirming the accuracy of the design procedure previously 570 described . Finally , it is very important to underline that the obtained flexural strength is 571 greater th an the plastic resistance of the connected beam, so that practically full -strength 572 connections are obtained without providing any damage to the beam ends. This is highlighted 573 in Fig.19 reporting with dashed lines the bending moment corresponding to the nom inal 574 beam resistance. In particular , it is worth to observe that even though no damage was 575 detected in the elements of the connection or in the member s, the nominal bending resistance 576 of the beam evaluated at the level of the column face was equated or exceeded. This shows 577 that with the proposed c onnection it is possible to obtain a wide and stable hysteretic 578 behaviour with negligible damage, carrying the same forces that the beam would normally 579 carry with a full -strength connection. 580 -27- 581 582 Figure 19 : Moment -rotation curves of tested joints 583 584 6. CONCLUSIONS 585 The results of an experimental program devoted to evaluat e the possibility to equip steel 586 joints with friction dampers realized with pads coated with thermally sprayed alumin ium 587 have been presented. The main outcomes of the experimental programs can be summarized 588 as follows: 589 • The frictional interface “ thermally sprayed alumin ium on steel” is characterized b y 590 good values of the kinematic friction coefficient. In fact, compared to other me tallic or 591 -28- rubber materials tested in a past experimental program under the same loading 592 conditions it is able to exhibit greater values of the friction coefficient; 593 • The thermally sprayed alumin ium has, among the various advantages, the benefit to be 594 an economic material compared to rubbers or other metals typically used for 595 tribological applications such as brass; 596 • The friction coefficient of the interface decrease s as far as the pressure applied to the 597 interface increases. Therefore, in order to develop high values of the sliding force it is 598 advisable to limit the pressure applied on the frictional material , as an exam ple by 599 increasing the number of the bolts thus allowing the reduction of their tightening 600 torque ; 601 • The results of the experimental program on the friction damper have shown that a 602 good balance between the initial value of the friction coefficient and the degradation of 603 the sliding force during a cyclic loading history can be reached with a thickness of 604 alumin ium coating equal to 300 m; 605 • As the bolt preloading has not been directly measured and monitored throughout the 606 tests, it is not possible to quantify in detail the actual value of the friction coefficient 607 and its evolution during the tests. Notwithstanding, the overall sliding resistance of the 608 damper is accurate as directly measured in the components ’ tests and derived as the 609 ratio between the sliding flexural resistance and the lever arm in the tests on beam -to- 610 column connections. This force is directly affected by the tighten ing torque; 611 • For the above reasons, new experimental tests are currently in progress where the bolt 612 preloading is continuously measure d by means of annular load cells . 613 614 ACKNOWLEDGEMENTS 615 The research activity herein presented has been supported by the European Community by 616 research grant RFSR -CT-2015-00022. The support of the European Commission within RFCS 617 Research & Innovation is gratefully acknowledged. 618 619