Beam -to-column joints equipped with friction dampers are promising solutions to improve the performance of steel moment resisting frames due to the possibility to guarantee large dissipation capacity limiting the structural damage under severe seismic cond itions. In this paper, the experimental tests and the numerical simulations of two types of joints are shown and discussed with the aim of developing pre -qualified configurations. The friction dampers are designed to be easily removable from both the lower beam flange and the column face by means of bolted connections. The devices are composed of a stack of steel plates conceived to assure symmetrical friction . The friction surface is set in vertical direction in first case and in horizontal direction in th e second type. The experimental tests confirmed the effectiveness of both examined joints and the finite element analyses allowed characterizing their local response, thus providing additional insights to improve the design requirements . Keywords: steel beam-to-column joints ; seismic response ; Connections ; friction damper; cyclic tests ; finite element simulations 1 Department of Civil Engineering, University of Salerno, Fisciano, Italy; mlatour@unisa.it 2 Department of Structures for Engineering and Architecture, University of Naples “Federico II”, Via Forno Vecchio 36, 80134 Naples, Italy; mdaniel@unina.it , Corresponding Author 3 Department of Structures for Engine ering and Architecture, University of Naples “Federico II”, Via Forno Vecchio 36, 80134 Naples, Italy; mariana.zimbru@unina.it 4 Department of Civil Engineering, University of Salerno, Fisciano, Italy; g.rizzano@unisa.it 5 Department of Civil Engineering, University of Salerno, Fisciano, Italy; v.piluso@unisa.it 6 Department of Struc tures for Engineering and Architecture, University of Naples “Federico II”, Via Forno Vecchio 36, 80134 Naples, Italy; landolfo@unina.it 1 Introduction The design criteria currently implemented by seismic codes are based on the philosophy of hierarchy of resistances, which aims to guarantee overall ductile and dissipative behavior by enforcing the plastic deformation , namely the damage , into specific ductile zones of the structure. However, e conomic and social reasons have recently pushed researchers and designers towards systems that can resist severe ground motions with low or without structural damage [1]. Traditional seismic resisting systems widely adopted for steel buildings (e.g. moment resisting frames, concentrically and eccentrically braced frames) entail a dissipation mechanism based on plastic deformation s of several structural elements , which may correspond large repairing costs in the aftermath of a seismic event . Therefore, t he idea of low or free f rom damage structures has become very appealing in the last decades [2-4]. The use of friction connections is a viable and promising strateg y to achieve this ob jective for steel structures [5-24]. In the framework of Eurocodes , this type of connections can be classified as partial strength according to EN1993:1 -8 [25] because their design resistance should be lower than the strength of the connected members to prevent any damage into the primary structural members . EN1998 -1 [26] allows the use of partial strength connections provided that their rotation capacity is properly demonstrated . In the case of conventional partial strength joints the ductility can be designed by imposing local hierarchy rules among the components constit uting the joints [27-31] and verified by means of pre -qualification tests [32- 34]. Moment resisting friction connections are conceived to develop the dissipation mechanism by means of the relative slip into ad -hoc devices located between the lower beam fl ange and the outer cap plate connected to the column flange , while the upper flange of the beam is connected to a plate either bolted or welded to the column. The cover plate connecting the upper beam flange may be subjected to some moderate plastic bendin g deformations to accommodate the joint rotation following the sliding of the device, thus enforcing the formation of an ideal center of rotation that prevent s the damage of the slab. To increase the moment capacity, friction devices can be also adopted for the beam web. In addition, the resistance of the joint can be modulated keeping the same assembly but varying the friction resistance that changes with the clamping force used for the bolts. Indeed , the friction device is composed of a stack of steel plates that are clamped together by means of tightened high strength bolts , which are inserted in the slotted holes of the plates to allow the relative sliding. The non -linear response of these connect ions depends on the type of friction mechanism , which can be either asymmetric or symmetric . The asymmetric friction connection (AFC) has been thoroughly investigated [7-10] and even successfully implemented in recent practice [ 1]. However, the bolts that clamp the friction pad of AFCs can experience yielding due to large bending moment, shear and axial force interaction, which can induce clamping loss of the bolts and consequently pinching and loss of strength of the connections. On the contrary, thes e phenomena can be mitigated with the use of a symmetric friction connection (SFC) [16 -22]. Latour et al. [2 3] recently carried out an experimental study on SFCs with the friction damper applied by means of an additional haunch welded to the lower flange o f the beam. The friction pad was located at the interface between the haunch and L -stubs that are connected to the column flange. It was composed by an 8mm S275JR steel plate (steel hardness 211HBW, sand blasted surface) coated with a thin layer of sprayed aluminum to improve the friction resistance with low cost of the raw material, as also shown by [ 35]. The tests showed that this solution is very effective, because the allowable flexural strength of the connection at column face is greater than the plast ic resistance of the connected beam . Hence, full-strength connections can be also obtained without requiring any damage to the beam. In addition, the use of a haunch increases the stiffness of the connection, since its internal lever arm is larger than the beam cross section depth , thus obtaining rigid connections prior the sliding of the device. These features are very important from design point of view, because the structural models adopted for the seismic analysis do not need to account for the deformab ility of the joints. It is also worth noting that t he geometry of these connections is similar with a split tee connection with the particularity of the friction pads needed in order to ensure specific friction properties. The resistance of the connection is dictated by the friction damper but, once the design resistance is established, the design of such connection can be carried out entirely according to EC3 Part 1 -8 [25]. In addition, the constructional costs for such connection are marginally different compared to traditional steel bolted connections. The economic advantages of the MRFs equipped with friction connections include also the limited extent of the damage that is localized at the level of the joint components, this further simplifying the rehabilitation work in the earthquake aftermath [15,16]. However, after severe seismic events the friction pad should be substituted, and the surface of the haunch should be treated to rest ore its initial roughness and to remove the residual portions of the friction layer scraped out the pad. These types of interventions can arise some operational difficulties especially concerning the tightening of the bolts that clamp the friction device. Indeed, as shown by [ 21, 23] it is crucial to control and guarantee the level of bolt pre -loading to ensure the design value of friction resistance. Indeed, if the clamping force is excessive ly large , the corresponding strength of the connection can be larger than the resistance of the adjacent members. On the contrary, lower preloading may either anticipate the sliding of the connections under serviceability non -seismic loading or weaken the global structural capacity under the design earthquake that ma y induce disproportionate rotation demand of the connections. A viable solution to solve these issues can be the use of removable friction dampers that can be easily detached from both the lower beam flange and the column face by means of bolted connectio ns, thus simplifying the reparability of the friction device. Indeed, if the whole friction damper is conceived as a demountable kit containing both the friction pad and the relevant steel supports, this option allow s tightening the bolts in the shop with the reliable control of the applied torque. In addition, the friction kit can be entirely substituted in the aftermath of severe earthquake without the need to perform superficial treatments of the beam flange on site. These considerations motivated the r esearch activity presented in this paper , which was devoted at developing two types of connections with detachable friction dampers . To achieve this objective both experimental cyclic tests and finite element simulations have been carried out . The paper is organized in three main parts as follows: i) the design criteria of the proposed joints are presented in the first part; ii) the experimental campaign is described in the second part; iii) the finite element simulations and the characterization of the loc al response of the joints are discussed in the third part. 2 Design criteria of joints 2.1. Features of the investigated joints with removable friction dampers The examined joints are characterized by double split -T connections, where the bottom tee element is replaced by detachable friction dampers, as depicted in Figure 1. The main mechanical difference of the two investigated types of devices is the direction of the friction plan that is horizontal in the case of bolted haunch (hereinafter also identified as “configuration 1”) in Fig. 1a, and vertical in the cas e of the bolted rib plate (hereinafter also identified as “configuration 2”) in Fig. 1b. In both cases the friction pads are made of stainless steel covered by a special type of coating selected based on former studies carried out at University of Salerno [19,20]. a) Configuration 1 b) Configuration 2 c) Components of the friction damper for configuration 1 c) Components of the friction damper for configuration 2 Figure 1 Typologies and features of the investigated j oint configuration s The friction kit of the configuration 1 is made of four main components (see Figure 1c), namely the L - stubs, the friction pads, the pre -loadable grade 10.9 HV bolts and the removable haunch, which is detailed with slotted holes on the lower side to facilitate the sliding at th e interface with the friction pad, while clearance fit bolts are considered for the upper part of the haunch to prevent relative its sliding as respect to the lower beam flange. The L -stubs are welded built -up and have the role of (i) transferring the bolt forces, (ii) clamping together the friction pads and the haunch, and (iii) connecting the assembly to the column flange. The friction damper of configuration 2 (see Figure 1d) has vertical sliding surfaces. Hence, to allow the relative slip both the mobile part (i.e. the rib plate) and the fixed parts (i.e. L -stubs and friction pads) are detailed with slotted holes. 2.2 Design criteria of the friction connections The moment resistance of the friction connections (Mslip,Rd) is assumed equal to the design moment at column face (MEd), which can be set equal to either the value obtained from the structural analysis under seismic condition (MEd,E) or the fac tored resistance of the beam cross section ( Mpl,b,Rd) projected at column face and scaled down by the factor Ωμ that accounts for the expected overstrength developable by the friction device in order to avoid damage in the beam at the haunch/rib tip . In the first case the beam -to- column assembly should have moment resistance larger than the moment induced by non -seismic loads at ultimate limit state of non -seismic conditions and it behave as rigid -partial strength joint, while rigid -full strength in the second case . In this study the specimens were designed according to the second approach in order to exploit the maximum resistance of the assembly . Generally speaking, w hichever is the design option, set as MEd the design bending moment of the connectio n, the required slip resistance of the device can be obtained as follows: ,=Ed slip req tMFh (1) Where ht is the lever arm of the connection, namely the distance between the center of rotation and the axis of sliding. In the present study, in order to exploit the larger flexural strength of the connection keeping elastic the spanning beam, the design moment demand of the connection was set equal to the factored plastic resistance of the beam projected at column face, namely as following: Ed pl b Rd h hM M S V,,()+ = (2) ( ) 2h pl b Rd h Ed GLV VM, ,,/+ = Where Sh is the distance from the tip of the haunch/rib plate to the column face and, being Lh the free span length of the beam from tip to tip of the haunches and VEd,G is the shear force due to gravity loads. In the cases of the experimental specimens of this study VEd,G was set equal to zero. The number of friction interfaces (ns), the mechanical features of the characteristic dynamic friction material (μdyn,5% ), the number (nb) of bolts and their diameter (d) were fixed a -priori . Subsequently, b y imposing that the effective slip force Fslip,eff has to be larger or equal to the required value (see Eq. 1) it can be easily determine the necessary level of bolt preloading (Nb) as follows: , , ,5% , ,5%slip req slip Rd dyn s b b slip req b dyn s bFF n n N F Nnn=     →  (3) In order to limit the relaxation of pre -loading due to creep phenomena and to prevent the yielding of the bolt shank under bending , on the basis of previous experimental studies [ 21, 35] the optimal values of the clamping force should range within 30% to 60% of preloading force ,pCF recommended by EN1993:1 -8 [25]. The design values for Nb are given in Table 5 while the friction material properties are given in Table 7. The non-yielding compone nts of the joints are designed according to EN1993:1 -8 [25] to resist the slip force and the associated moment magnified by the overstrength factor Ωμ , which is defined as follows: .95% .95% .5% .5%1.97 2.0  =  = st b dyn bN N (4) Where μdyn,5% , Nb,5% are the lower -bound values of the dynamic friction coefficient and tightening force, respectively; μst,95%, Nb,95% are the upper -bound values of the static friction coefficient and tightening force, respectively. The values of both static and dynamic friction coef ficients as well as the tightening forces were derived on the basis of previous experimental studies [ 18, 19, 20, 21, 35 , 45] which addressed aspects related to the friction properties of the coating material and the covered surfaces , the preloading level, creep phenomena, loss of preloading and high velocity load application. On the basis of these studies , the selected friction material is assumed to have predictable and stable friction properties under very severe cumulated displace ment histories. However, further studies should be carried out to investigate any aging deterioration along the design life . Another important check that should be considered is the shear resistance of both the T -stub and L -Stubs at column face . The shear resistance of these connections is evaluated according to EN1993 1 -8 [25], while the design shear force VEd is given by Eq. (5): Where it can be noted that the design shear force is estimated as the ratio between twice the expected strength of the friction joint (MCD) and the distance between the column face connections ( L). In addition, both the T -stub and the L -stubs connections have to be checked to resist the full value of VEd,. The explanations and the relevant evidence of this recommendation are given hereinafter in Section 4.2. 3 Experimental campaign 3.1 Test Setup The experimental campaign consisted of 4 tests on friction joints that were performed in the laboratory of the Department of Civil Engineering at University of Salerno . Figure 2 Experimental setup The test setup is shown in Fig. 2. The column s were pinned in the bending plane of the joint with a cylindrical hinge at one end and a roller at the opposite tip . The beam was laterally restrained by means of a braced frame to prevent its lateral -torsional buckling. Two MTS 243 actuators were used for the tests. The top actuator (load capacity of ±250 kN and stroke range equal to ±500 mm) was used to apply t he displacement history at the beam tip that was set according to loading protocol recommended by AISC341 [36]. The second actuator (load capacity of ±1000 kN and stroke range equal to ±125 mm) was used to apply a constant compression force equal to 30% of the column squash load that was kept constant throughout the duration of each test . 3.2 Investigated joints and monitored parameters As previously discussed, two types of friction dampers were investigated, and their main geometrical features are depicted in Fig. 3a -g. In addition, two beam -column assemblies per joint type were tested to cover the cases of both small (beam IPE270 – column HE 220M) and large (beam IPE450 – column HE 500B) scale s tructures. In all tested joints the bolts of the friction devices were equipped with special type of disc springs (DS) washer s to limit the loss of preloading under cumulated slip , which are non -flat washer with a slight conical shape and are commonly used to solve vibration, thermal expansion, relaxation and bolt creep problems. The DS were stacked both in parallel and in series to increase the strength and the deformability, respectively as shown in Fig. 3. This arrangement was established on the basis of the findings 2CD EdMVL= (5) CD EdMM= of a former experimental campaign on lap -shear connections [18, 19, 35]. In addition, preloadable HV bolts [37] were adopted to clamp the friction surfaces. a) T-stub b) L-stub (Confiuration 1) c) Haunch (Confiuration 1) d) L-stub (Confiuration 2) e) Rib (Confiuration 2) f) Friction pads (Confi guration 1) g) Friction pads (Confi guration 2) h) Adopted stacking of Disk Springs Figure 3 Geometrical details of the tested friction devices Table 1 Specimen geometrical configuration Specimen Connection type Beam Column Length of the Beam “L” Length of the column “H” Internal level arm of the connection “z” Thickness of beam web stiffener “tstiff, b eam” Thickness of column continutiy plate “tstiff, c olumn” mm mm Mm mm mm FD-1-1 1 IPE270 HE 220M 1556 2700 478 10 10 FD -1-2 1 IPE450 HE 500B 2906 2700 710 15 20 FD -2-1 2 IPE270 HE 220M 1556 2700 458 10 10 FD -2-2 2 IPE450 HE 500B 2906 2700 710 15 20 Table 2 T-stub geometry details Flange Web Specimen h b tfl Bolts e1 p1 e2 p2 l tw Bolts e1 p1 e2 p2 mm mm mm mm mm mm mm mm mm mm mm mm mm FD-1-1 187 200 20 M20 43 101 48.3 103.4 235 15 M18 30 50 58.3 83.4 FD-1-2 270 300 30 M30 70 130 75 150 410 20 M24 55 65 95.6 108.8 FD-2-1 195 200 20 M20 45 105 38.3 123.4 260 15 M18 30 50 58.3 83.4 FD-2-2 270 300 30 M30 70 130 75 150 485 20 M24 50 65 95.5 109 Table 3 L-stub geometry details Flange Web Specimen h bL bLL tfl Bolts e1L e1LL e2 p2 L tw Bolts e1 p1 e2 p2 mm mm mm mm mm mm mm mm mm mm mm mm mm mm FD-1-1 98 200 85 20 M20 40 43 42.5 115 219 20 M20 62 70 42.5 115 FD-1-2 145 300 135 30 M30 60 60 67.5 165 360 30 M20 85 70 67.5 165 FD-2-1 220 97.5 20 M20 60 100 40.5 260 20 M20 70 70 40 160 FD-2-2 330 132 30 M30 90 150 57 366 30 M20 83 70 65 200 Table 4 Haunch geometry details Specimen Ltop Lbot tfl tw Bolts e1 p1 e2 p2 mm mm mm mm Mm mm mm mm FD-1-1 419 234 15 10 M18 35.5 67 25.8 83.4 FD-1-2 590 365 20 15 M24 40.4 81 40.5 109 FD-2-1 520 245 15 15 M20 72.1 72.1 25.8 83.4 FD-2-2 756 356 20 20 M24 91.7 91.7 40.5 109 Table 5 Friction pads geometry details Specimen l b b t Bolts Disc spring Nb e1 p1 e2 p2 mm mm mm mm kN mm mm mm mm FD-1-1 194 200 85 8 M20 w/ w/o 72 62 70 42.5 115 FD-1-2 310 300 135 8 M20 w/ w/o 87 85 70 67.5 165 FD-2-1 210 220 8 M20 w/ w/o 78 70 70 40 160 FD-2-2 306 330 8 M20 w/ w/o 96 83 70 65 200 The overall geometrical features of the tested beam -to-column assemblies are depicted in Fig. 3 and the corresponding values are also reported in Table 1 to 5. A label code is used to identify each assembly as follows: ‘FD’ stands for Friction Device, the first number is related to the joint configuration (1 or 2) and the second number refers the size of the assembly (1 for small or 2 for large). The a dditional term “DS” identified the use of disc springs. As example, FD -1-1-DS corresponds to the small assembly equipped with the device configuration 1 (i.e. horizontal friction surface) and disc springs. Both global and local response parameters were mon itored during the tests. At global level, the displacement s at the tip of the beam and the relevant reaction force s of both actuator s were measured (see Fig. 4a) . Namely, the bending moment in column axis was evaluated based on the reaction force in the actuator ( F) and the corresponding lever arm ( L – the distance between the actuator and the column axis). The displacement ( u) at level of the actuator applying the displacement history is used to obtain the chord rotation (θchord) by dividing it with L. Regarding the local response parameters, the relative displacements (δ) among the components of the connections were monitored by means of LVDTs and the forces into the bolts were measured throughout the experiments by means of torsional loading cells. The relative sliding of the friction damper with respect to the fixed parts (L -stubs) were measured by means of the LVDT -07 (for the configuration 1) and LVDT - 02 and LVDT -07 (for the configuration 2) , as depicted in Figure 5. The obtained values were used to evaluate the friction damper rotation (θdevice), which is obtained by dividing the displacement δ with the connection lever arm z measured from the mid -thickness of the T -stub web to the geometrical center of the friction device as detailed in Figure 4 . In addition , LVDT -03 was used to evaluate the opening of the cap Tee stub connection, while LVDT -04 (for the configuration 1) and LVDT -05 (for the configuration 2) were used to monitor the gap opening of the lower L -stub connections. a) b) c) Figure 4 Monitored parameters: global (a) and local response (b,c) a) Joint FD -1-1 b) Joint FD -2-1 Figure 5 Layout of displacement transducers to measure the local response of joints 3.3 Test results All performed tests showed an overall satisfactory response with stable hysteretic behavior and similar features , as depicted in Figure 6. Indeed , both types of connections performed as rigid up to the static friction resistance of the devices. Following the activation of the sl iding , a loss of strength was observed but negligible stick -slip phenomena were observed under load reversal. Both friction assemblies lead to an asymmetric response of the joint, however the difference between the sagging and hogging resistance is larger in the case of the configuration 1 (i.e. horizontal friction surface), which also experienced the more pronounced degradation of the cyclic resistance, especially for the large assembly FD -1-2-DS. The different flexural resistance experienced under sagging and hogging can be explained by analyzing the local deformation demand in the upper T -Stub (which connects the top flange of the beam to the column) Center of rotation z Fslip Center of rotation z Fslip and the lower L -Stubs (which connect the device to the column). The upper Tee was the same at the same beam-column assembly of both joint types, thus the differences can be explained by considering the different deformability of the L -Stub arrangements that differ with the friction connection configurations. As depicted in Figure 7, the deformation of the Tee has values in the same ranges for the 2 pairs of tested joint assemblies i.e. the small assemblies (i.e. FD -1-1-DS and FD -2-1-DS) and the large assembl ies (i.e. FD-1-2-DS and FD -2-2-DS) present T -stub flange opening of maximum 0.5mm and 1.0mm, respectively. On the other hand, Figure 8 shows that the L -stub flange opening is substantially larger in the joint assemblies detailed with the horizontal friction damper configuration (i.e. FD -1-1-DS and FD -1-2-DS) than the values measured for the joints with the vertical damper (i.e. FD -2-1-DS and FD -2-2-DS). It is worth noting that the difference is terms of gap opening corresponds to the difference in terms of flexural strength under sagging loading. In line with the findings of a previous numerical study carried out by the Authors [3 8], the difference between sagging and hogging bending resistance for the first configuration can vary up to the 25%, while the second configuration has a lower difference (about 15%). a) Small assemblies b) Large assemblies Figure 6 Comparison between the response of the 2 device configurations a) FD -1-1-DS b) FD -1-2-DS . c) FD -2-1-DS d) FD -2-2-DS Figure 7 T-stub opening versus damper rotation -300-200-1000100200300 -0.06 -0.03 0.00 0.03 0.06Bending Moment [kNm] Chord Rotation [rad]FD-1-1-DS FD-2-1-DS -1000-800-600-400-20002004006008001000 -0.06 -0.03 0 0.03 0.06Bending Moment [kNm] Chord Rotation [rad]FD-1-2-DS FD-2-2-DS a) FD -1-1-DS b) FD -1-2-DS c) FD-2-1-DS d) FD -2-2-DS Figure 8 L-stub opening versus damper rotation As shown in Fig. 6, the larger beam -column assemblies exhibit ed the larger degradation of the strength for increasing levels of rotation. This performance was caused by two effects: first, the use of longer bolts (with the same diameter of those used for small er specimens ) and secondly, the bolt clamping was higher ( see Table 6 that reports the average tightening forces measured per specimen ). The former feature induced more flexibility in the bolts , which corresponded larger bending demand in the shank as well as rotation of head and nut a consequential reduction in clamping force that translates in to a reduc tion of the connection bending capacity. The second feature directly affect ed the threshold and the evolution of flexural strength of the connection, because assuming larger clamping level ( e.g. closer to the EC3 1 -8 recommended value) corresponds to reduce the strength margin of the bolt shank as resp ect to the plastic damage , which can be anticipated at lower slip displacements of the device. Furthermore, the horizontal and vertical damper configuration s induce different bearing forces on the clamped shank of bolts , which leads to different pattern of local plastic strain (see Figure 1 3 in the Section 4.2) and strength degradation when the rotation demand increases . The experimental tests showed negligible plastic deformation in the beam or the T and L -stubs up to 0.04rad of chord rotation. Increasing the imposed rotation up to 0.05 rad (which was set as the limit of the allowable displacement capacity of the friction device) the overall response of the joints was still satisfactory , as shown in Figure 9, without appreciable damage except for the unique case of FD -1-2-DS where the instability of the beam web occurred. This phenomenon occurred to activation of a strut transmission force betwee n the theoretical center of rotation around the Tee web connecting the upper beam flange and the device, which has been better clarified by means of finite element simulation hereinafter (see Section 4.2). However, it is worth noting that this phenomenon c an be influencing solely for deep beams and it can be easily prevented by using transverse web stiffeners at the tip of the beam. After the tests, it was observed that the friction pads underwent significant erosion of the superficial friction coatings an d signs of plastic deformation in the bolts were also spotted. Table 6 Pretension levels in the bolts of the tested specimens Model Nb [kN] Nb/Fp,c [%] FD-1-1 58 34% FD-1-2 98 57% FD-2-1 64 37% FD-2-2 105 61% -4,00-3,00-2,00-1,000,001,00 -0,100 -0,050 0,000 0,050 0,100L-stub opening [mm] Damper Rotation [rad] -4,00-3,00-2,00-1,000,001,00 -0,100 -0,050 0,000 0,050 0,100L-stub opening [mm] Damper Rotation [rad] -1,000,001,002,003,004,00 -0,100 -0,050 0,000 0,050 0,100L-stub opening [mm] Damper Rotation [rad] -1,000,001,002,003,004,00 -0,100 -0,050 0,000 0,050 0,100L-stubopening [mm] Damper Rotation [rad] a) FD-1-1-DS b) FD -1-2-DS c) FD -2-1-DS d) FD -2-2-DS Figure 9 Deformed shape of the t ested specimens at chord rotation equal to 0.05 rad 4 FEM 4.1 Modelling assumptions The experimental tests were used to validate the finite element (FE) models developed using Abaqus v 6.14 [39]. The quasi -static analyses were performed by employing the Dynamic Implicit solver. The geometrical characteristics of the experimental assemblies were replicated in the FE software by modelling solid parts meshed using the C3D8R finite element (an 8 -node linear brick with reduced integration). Both geometrical and mechanical nonlinearities were accounted for. The experimental tests on the large beam -to-column assemblies have suffered unexpected slippage in the connections between the column and the setup and therefore, the deformability source was taken into account in the numerical model. Figure 10 depicts the general model shape and the average material properties obtained experimen tally. The steel material properties were modelled based on the coupon tests performed in the laboratory as part of the experimental campaign , namely the yield stress was set equal to 380 MPa for beams, 427 MPa for columns and 443MPa for both L -stub and T -stubs . The elastic modulus was assumed equal to 210000MPa and Poisson’s coefficient equal to 0.3. The nonlinear branch of the constitutive law was implemented using a half cycle input method and assuming both nonlinear kinematic and isotropic plastic hard ening, as described by [ 30-33,40 ,41,42]. The bolts were modelled as shown by [ 43-44]. Meshed details Material properties Figure 10. Features of FE model The model parts in contact , such as the bolts and plates, were assigned with interactions modelling both the Normal Behavior to avoid overclosure (by means of the “Hard Contact” option) and Tangential behavio r to define the relative sliding (by employing the Coulomb friction law) [46]. In addition, to simulate the partial loss of the friction coefficient due to the smoothing of the superficial roughness of the friction pad, the temperature -dependent friction laws developed by the Authors [4 4] was used, thus friction coefficient decreases with the increase of temperature because of continuous sliding of plates. The reference friction properties are presented in Table 7. The simulations shown hereinafter were performed assuming the 5% dynamic percentile was used for the numerical simulations. Table 7 Friction material properties Friction coefficient 5% Percentile μ5% 95% Percentile μ95% Static 0.69 0.84 Dynamic 0.53 0.65 Since no plastic deformations are expected in the welded components, t ie constrain ts linking together surfaces in contact have been used to replicate in a simplified manner the presence of full penetration welds. The analyses were performed considering two loading steps: (i) bolt clamping and (ii) displacement history application. The boundary conditions of the joints have been accurately simulated to reproduce those used during the tests. In addition, the beam was lateral ly restrained with out -of-plane restraints located in the same sections of the experimental setup . The AISC 341 [ 36] loading protocol up to 5% interstorey drift ratio was applied at the beam end consistently with the test procedure . 4.2 Validations and discussion of results The adopted modelling assumptions effectively simulate both the global and the local response of the tested joints , as it can be observed in Fig. 1 1 and 1 2 respectively . Since the transition from the static to the dynamic friction was not modelled, t he response of the joint during the initial cycles are not accurately replicated . However, this inaccuracy disappears at increasing the number of cycles. a) FD-1-1-DS b) FD -1-2-DS c) FD -2-1-DS d) FD -2-2-DS Figure 11 Experimental vs numerical results in terms of Bending Moment – Connection rotation -300-200-1000100200300 -0.06 -0.03 0.00 0.03 0.06Bending Moment [kNm] Connection Rotation [rad]Experimental Numerical -1000-800-600-400-20002004006008001000 -0.06 -0.03 0.00 0.03 0.06Bending Moment [kNm] Connection Rotation [rad]Experimental Numerical -300-200-1000100200300 -0.06 -0.03 0.00 0.03 0.06Bending Moment [kNm] Connection Rotation [rad]Experimental Numerical -1000-800-600-400-20002004006008001000 -0.06 -0.03 0.00 0.03 0.06Bending Moment [kNm] Connection Rotation [rad]Experimental Numerical a) Experimental joint FD 1 -1 b) Experimental joint FD 2 -1 c) PEEQ distribution for model FD 1 -1 d) PEEQ distribution for model FD 2 -1 Figure 12 Experimental vs. numerical models after cyclic test up to 5% a) FD -1-1 b) FD -1-2 c) FD-2-1 d) FD -2-2 Figure 13 Equivalent plastic damage (PEEQ) in the damper bolts During the experimental campaign no damage was observed in the steel elements. However, the numerical analyses show some concentrations of slight plastic damage, depicted in terms of equivalent plastic strain (PEEQ) in Fig. 12, at the base of the web of the upper T -stub (where the center of rotation is located), and either at the bases of the L -stubs, at the top web -flange area of the beam underneat h the T -stub and in the shear bolts of the device. Furthermore, plastic deformations can be observed in the shanks of the bolts in the friction device. Indeed, the horizontal damper configuration induces shear type bending effects in the shanks with two be aring contacts in all bolts of the device. On the contrary, the bolts in the vertical damper have one bearing zone at mid -length of the clamped shank, which leads to larger local plastic strain (see Fig. 13). In addition, in this second case the bolts clos e to the column face do not exhibit plastic strains. A crucial aspect related to the design of the connections (T -stub and L -stub) at column face is the shear check, because these elements should guarantee the resistance due to combined tensile and shear forces to avoid premature failure. Figure 14 shows the distribution of shear forces at the level of the Tee and L -stubs as well as the total shear force in the section at the column face. In both tested friction device configuration s, the cumulated shear in the two components (i.e. the sum of the releva nt absolute values) is larger than the overall shear force (see Figure 14 a and c for type 1 and Figure 14 b and d for type 2) . In order to investigate the evolution of the shear force with the connection rotation, monotonic analyses under both hogging and sagging loading conditions were alternatively performed . The results presented in Figure 15 confirm the previous observations and offer insight into the magnitude of the shear t ransferred by each component. Indeed, for the investigated cases, the L -stubs transfer larger shear force compared to the T -stub. Configuration 1 joints are characterized by levels of maximum shear force transferred by the Tee of about 50% of the total she ar, while the L stub s reach values close to 100% of the total shear (Figure 15 a and c). However, while Configuration 1 components transfer shear of u p to maximum total shear, Configuration 2 assemblies ( Figure 15 b and d) exhibit the same behavior observed cyclically i.e. the shear in the components reach values larger than the total shear, with the shear transferred by the L -stubs reaching values almost 2 times larger than the total shear for rotation values close to 0.0 6 rad. a) FD -1-1 b) FD -1-1 c) FD -1-2 d) FD -2-2 Figure 14 Shear force transferred by different components a) b) Hogging c) d) Sagging Figure 15 Shear force at the connection face The differences of shear force distribution s between the 2 components is mainly due to the larger stiffness provided by the L -stubs in the vertical plane and the vertical sliding force component . Indeed, the transfer mechanism of shear force amongst the components (see Figure 16 ) is highly complex and configuration dependent. The type 1 joint (FD 1 -1-DS) is characterized by same sign shear forces transferred by the L - stubs and in a smaller measure by the T -stub. On the contrary, the T -stub of FD 2 -1-DS carries an opposite sign shear force, owing to the heightened level of shear force transferred by the L -stubs , in order to preserve the equilibrium at the column face. The hogging/sagging loading conditions lead to the same distribution of forces between the components for the same configuration , with a smoother transfer of the forces under positive bending . FD 1 -1-DS FD 2 -1-DS Figure 16 Distribution of shear in the connected elements at 0.04 rad damper rotation In the analyses performed on the models depicting the experimental tests it was observed that small concentration of damage is located in the base of the T -stub, the L -stubs (in the case of Configuration 1), a) Hogging b) Sagging the bolts of the friction damper and in a reduced extent in the beam (the web -flange junction at the beam end immediately below the T -stub and in the slotted holes at the end of the beam -haunch connection). These results are p resented in Figure 17 in terms of PEEQ (equivalent plastic strain) distribution on the large beam to column assemblies. As it can be observe d in the legend of PEEQ that are evaluated at the rotation equal to 0.05rad , the horizontal friction damper configuration leads to larger plastic deformations in the joint elements. As a matter of fact, this result can also be observed in terms of dissipated energy , presented in Figure 18. Indeed, the friction energy normalized with respect to the total dissipated energy for the first configuration is smaller compared to the second configuration and the opposite is true in terms of normalized plastic energy. Although the second damper configuration leads to lower plastic damage, it is worth noting that the plastic damage is limited for both configuration, maximum plastic energy dissipation is less than 5% of the total energy dissipated at 0.05 rad. a) FD -1-2 b) FD -2-2 Figure 17 PEEQ Distribution at the end of the cyclic analysis for large joint assemblies a) b) Figure 18 Normalised disipated energy a) Friction energy and b) Plastic energy 5 PARAMETRIC FE ANALYSIS 5.1 Generality The experimental tests confirmed the importance of the clamping force applied to the bolts into the device as well as the ke y role of the friction coefficient between the sliding surfaces. Therefore, the influence of the variation of these parameters on the response of the two tested joint configurations has been further investigated by means of finite element analyses based on the validated models described in Section 4. The examined parameters have been varied as follows : • In addition to the pre -loading force adopted in the tests, namely Nb, a value 50% smaller (0.5 Nb) and a value 50% larger (1.5 Nb) have been considered . It should be noted that in all cases 1.5 Nb is smaller than Fp,C (which is equal to 172kN for M20 gr.10.9 bolts). • Three values of the dynamic friction coefficient μ are considered , namely the 5% percentile ( μ5%), see Table 6 , the average value ( μavg) set equal to 0.59, and the 95% percentile ( μ95%), see Table 7. 5.2 Influence of c lamping force Figure 19 shows the comparison of the response curves for the four joints (i.e. the two joint configurations and two assemblies) and Table 8 and Table 9 report the numerically measured mechanical properties of the joints. The bending moments reported, M (+) and M (-), represent the bending moment at the occurrence of the sliding under sagging and hogging , respectively . Equations ( 6), (7) and ( 8) clarify the meaning of the mechanical parameter reported in the tables. ( ) ( ) 0.5 () ()1bb bNN NMM M++ + +− = + (6) ( ) ( ) 0.5 () ()1bb bNN NMM M−− − −− = + (7) ( ) ( ) ( / ) ( ) ()/MMMMM−+ + − − −−= (8) Where Γ(+) and Γ(-) represent the variation of the hogging and sagging bending moment capacity, respectively, considering alternatively the change in the clamping force from the design value Nb to 0.5 Nb and 1.5 Nb; M (+) and M (-) are the sagging and hogging bending moments. The subscripts depict the analysis from which the bending moment is taken, e.g. with clamping force equal to either 1.5 Nb or 0.5 Nb; ΔM (+/- )/M (-) represents the difference between the hogging and sagging bend ing moment for each respective analysis (considering the three values for Nb). As expected, the variation of the bending moment is proportional with the bolt pre -tension, although differently under both sagging and hogging. As reported in Table 8 and Table 9, this difference is strictly related to the joint configuration and it is constant with the beam depths, clamping force, or friction coefficient. The difference is about 25% for Configuration 1 and 15% for configuration 2. Further observation that can be made based on Figure 19 is that the post -yield response of joint configuration 1 differ s with the size of the beam -to-column assembly and with the level of preloading (relative to the maximum preloading force). In particular, the joint with shallow beam and lower relative preloading exhibits hardening (i.e. positive post -yield stiffness), wh ile the joint with deeper beam and higher relative preloading shows softening (i.e. negative post -yield stiffness), the latter is more evident for the lower values of clamping force. These phenomena are more pronounced under hogging bending moment. The sec ond configuration exhibits a more linear behaviour in both examined assemblies. The stiffness of the joint is not affected by the variation of clamping force, since it is determined by the stiffness of the other components of the joint (the connection at c olumn face, the column web panel, etc.). Assembly 1 Assembly 2 Confguration 1 Configuration 2 Figure 19 Influence of the clamping force on the bending moment capacity Table 8 Bending moments for model FD 1 -2-DS considering the variation of clamping force Clamping force M (+) M (-) Γ (+) Γ (-) ΔM (+/-)/M (-) [kNm] [kNm] [-] [-] [-] Nb 453 602 - - 25% 0.5N b 230 298 51% 50% 23% 1.5N b 690 902 152% 150% 24% Table 9 Bending moments for model FD 2 -2-DS considering the variation of clamping force Clamping force M (+) M (-) Γ (+) Γ (-) ΔM (+/-)/M (-) [kNm] [kNm] [-] [-] [-] Nb 484 564 - - 14% 0.5N b 250 290 52% 51% 14% 1.5N b 714 838 148% 149% 15% 5.3 Influence of Friction Coefficient Figure 20 depicts the numerical curves in terms of bending moment vs. chord rotation. It is possible to observe that the higher percentile of the friction coefficient values the larger is the joint capacity. This observation confirms the need to account for the variability of the friction properties of the friction pads to design the non -yielding structural members. Similar hardening/softening behaviour can be observed for both joint configuration an d, additionally, the response curves seem scaled proportionality with the friction coefficient. Table 10 and Table 11 depict the variation of the bending capacity of the FD -1-2-DS and FD -2-2-DS models analysed with larger values of friction coefficient ( μavg and μ95%) with respect to the design value (μ5%) under hogging (M (-)) and sagging ( M (+)) loading conditions. The variation in the case of FD -1-2-DS differs with respect to the variation of the friction coefficient. In particular, a larger increase of bending moment can be observed for the same i ncrease of friction coefficient. On the other side, the analyses of model FD -2-2-DS in Table 11 show a closer dependency of the bending capacity with t he friction property randomness. Table 10 Bending moments for model FD 1 -2 considering the friction coefficient variation Friction coefficient Δμ M (+) M (-) Γ (+) Γ (-) ΔM (+/-)/M (-) [-] [kNm] [kNm] [-] [-] [-] μ5% - 446 593 - - 25% μavg 110% 521 670 117% 113% 22% μ95% 117% 535 733 120% 124% 27% Table 11 Bending moments for model FD 2 -2 considering the friction coefficient variation Friction coefficient Δμ M (+) M (-) Γ(+) Γ (-) ΔM (+/-)/M (-) [-] [kNm] [kNm] [-] [-] [-] μ5% - 484 564 - - 14% μavg 110% 529 627 109% 111% 16% μ95% 117% 568 679 117% 120% 16% The parameter Δ M (+/-)/M (-), evaluated also for this set of analyses confirms the previous observation regarding the relation between the damper’s configuration and the different response under sagging and hogging conditions (values ranging around 25% for configuration 1 and 15% for configuration 2). Assembly 1 Assembly 2 Configuration 1 Configuration 2 Figure 20 The influence o f the friction coefficient on the bending moment resistance 6 Conclusive Remarks The seismic response of steel beam -to-column assemblies equipped with two types of friction dampers has been investigated by means of both experimental tests and finite element simulations . Based on the obtained outcomes, the following remarks can be drawn : • Both types of friction joints provided satisfactory overall performance with stable and predictable hysteretic response, as well as preventing from damage the non -yielding members . However, non - symmetry response under sagging and hogging was observed. • The joint configuration dictates the level of the response symmetry under sagging and hogging bending. The configuration with vertical friction surface exhibited slightly better response, showing lower degradation under cumulated rotation demand with smaller difference between hogging and sagging resistance. Indeed, the different resistance under sagging and hogging conditions ranges abou t 25% for the configuration with horizontal friction surface and 15% for that with vertical friction surface. • The FE models accurately predict the response of experimental tests . The models allowed characterizing the local response of the joints, which ex hibit some plastic deformations in the bolts and in the connecting L -stubs and T -Stubs. The FE analyses also showed that the joints equipped with the vertical friction damper exhibit plastic deformations lower than the corresponding device with horizontal friction surface. • The variation of the bending capacity of both joint configurations is directly proportional with the bolt pretension force. Therefore, the bolt tightening process needs to be very well controlled because either larger or smaller tightenin g forces can impair the proper dissipative mechanisms. Indeed, the upper bound values lead to the development of larger forces in the damper, situation that hinders the hierarchy in the joint, while lower clamping forces can lead to sliding in the damper under serviceability conditions. • The randomness of the friction properties has to be as much as possible mitigated and accounted for in the design phase, because this variability can inflict in the joint response and, consequently, the global behavior of the structure. 7 ACKNOWLEDGEMENTS The research activity herein presented has been supported by the European Community by research grant RFSR -CT-2015 -00022. The support of the European Commission within RFCS Research & Innovation is gratefully acknowledged. 8