Yielding in steel structures dissipates seismic energy, which results in residual deformations. Conventional lateral -resisting systems provide appropriate seismic safety. However, these systems undergo severe structural damage when exposed to strong ground motions. Seismic residual drifts make it difficult or impossible to fix damaged structures. Therefore, systems that can reduce seismic residual deformations are required. Superelastic shape memory alloys (SMAs) can withstand significant deformations and r ecover all plastic deformations upon unloading. They can significantly lessen seismic residual deformations in steel structures, making post -seismic retrofitting easier. This paper will compare the use of a superelastic Ni -Ti alloy and a Fe - based SMA in tw o endplate connections of NIST benchmark buildings (NIST Technical Note 1863 -3). In addition, some features that lead to decreased residual deformation in these types of joints will be investigated. Keywords (Times New Roman bold 11 pt.) : Shape memory alloys, Superelastisity, Finite element method, self-centering , Endplate connection, Residual deformation . 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 2 Introduction Until the 1990s, Steel moment resisting frames with fully restrained welded beam -to-column connections (such as joints with welded beam flange) had long been accepted as a suitable structural solution against seismic loadings. The 1994 Northridge and 1995 Kobe earthquakes revealed the deficiencies in these connections. Such connections, which at the time represented the majority of connec tions, frequently failed due to weld fracture. Later, a study known as the SAC Project was started to look at the causes of the extensive weld fracture and, as a result, make recommendations for better connection specifications for steel structures against earthquakes (FEMA 2000). This project provided significant insight into the use of bolted steel connections in seismic applicat ions [1, 2 ], and it was found that when properly designed and detailed, bolted connections would provide high seismic performanc e with adequate strength, ductility, and stiffness for seismic performance. Relevant test results demonstrated that a "strong -connection - weak -beam" theory was the preferred design method, allowing plastic hinges to be developed over the beam portion close to the connection region [3-5]. Additionally, it was discovered that more flexible semi-rigid connections could be used as a substitute to provide adequate plastic rotations and reliable seismic perfo rmances [1 ]. The typical design approaches for seismic r esistant connections generally result in post -earthquake deformations in the beams (strong connections) or in the connections (partial strength connections), which are expensive and problematic to repair. Researchers are developing novel design strategies to reduce seismic residual deformations. For instance, to provide a recentering mechanism after an earthquake, special post-tensioned bars were installed in connecti ons [6, 7]. Shape memory alloys (SMAs) have also received much attention due to their self-centering and damping characteristics, and connections with shape memory alloys (SMAs) have lately appeared as a promising substitute for reaching the self -centering purpose [8]. The SMAs made of nickel -titanium (Ni - Ti) received the most extensive researc h [9]. Shape Memory Effect (SME) and superelasticity (SE) are the two main characteristics of SMAs. SME refers to a material's potential to recover from significant mechanically produced strains by gradually raising its temperature. SE refers to a material 's capacity to sustain large inelastic strains and recover its initial shape after load removal. Furthermore, the base isolations and dampers used as passive control devices, which have concerns with aging, durability, and maintenance, may no longer be nec essary due to SMA's high corrosion resistance performance (equal to stainless steel). An external beam -column connection using martensite SMA rods was put to the test by Ocel et al. [ 10]. The connection demonstrated excellent energy dissipation, significant ductility, and no strength degradation up to a 4% drift level. Besides, When the SMA tendons were heated, the connection recovered 76% of the previou s drift. DesRoches et al. [1 1] investigated SMA bars and wires, particularly SMAs with large cross -section s, to determine cyclic properties and their potential for seismic design and retrofit purposes. In this study, the effects of bar size, loading frequency, and yield stress on damping, recentering, and recentering capability was conducted to assess their fe asibility for utilization in seismic applica tions. DesRoches et al. [1 2] and Ellingwood et al. [1 3] used SMA bars in novel beam -to-column connections to evaluate the seismic performance of SMRFs incorporating these connections. The recentering characterist ic and high energy dissipation were produced using superelastic SMA connections and martensite SMA connections, respectively. Nonlinear time history studies were conducted to determine how these connections affected peak and residual inter -story drifts. An extended endplate connection using SMA bolts was suggested by Ma et al . [14]. The superelastic SMA bolts were used for this type of connection instead of the conventional high strength (HS) bolts. The deformation of SMA bolts accomplished the energy dissipation and ductility requirements. Thus, the superelastic hinge could develop inside the connection while the adjacent structural members (endplate, beam, and column) are mostly within the elastic range. This design approach's innovative benefit is reducin g the need for structural member repairs after an earthquake, which will significantly lower the cost of retrofit work. An experimental quasi -static analysis was performed by Ma and Yam [15] to look into the seismic performance of endplate connections that were equipped with eight long SMA bolts. Nearly all of the drift was recovered by unloading, plastic deformation was restricted to SMA bolts with relatively large capacities, and all beams and columns stayed within the elastic ra nge. Speicher et al. [1 6] experimentally evaluated four beams to column connections incorporating 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 3 superelastic Ni -Ti tendons, martensite Ni -Ti tendons, and a combination of Ni -Ti and Aluminum tendons under cyclic loading. After a maximum drift of 5%, both Ni -Ti tendons demonstrated a significant level of self -centering, with residual deformations measured between 0.5 and 1%. Extended endplate connections with superelastic bolts in an MRF were analyzed experimentally and numerically by Fang et al. [ 17, 18]. SMA connections exhibited high recentering ability and proper energy dissipation. Slender (long with a small diameter) SMA bolts represented more ductile connections. Farmani and Ghassemieh [ 19] numerically studied two innovative types of beam -column connections. Including SMA tend ons and shear - resisting elements like shear tabs and web hourglass steel pins into the connection improved the connection's energy dissipation, moment capacity, and shear resistance. However, if these components had been improperly designed, they would hav e compromised the connection's self -centering, stiffness, and energy dissipation. Sultana and Youssef [ 20] investigated where the SMA connections should be installed in various MRFs to minimize residual drifts after a seismic event while considering the ma ximum inter -story drift during an earthquake. Roshanfekr Rad et al. [ 21] probabilistically studied the seismic performance of two typical lateral load -resisting systems, including MRF and EBF, which incorporated superelastic SMA bolts in their beam -column connections and link-beam connections, respectively. Effects of SMA connections on peak inter -story drifts and residual drifts were investigated. Also, comparisons between conventional and SMA connections in terms of life safety performance and repair cost were performed. Although employing Ni -Ti SMAs to achieve structural self -centering after earthquakes has enormous potential, SMA components are expensive. A significant effort has been made to produce other SMAs, such as iron -based SMAs, which are appropr iate for use in large -scale civil engineering applications (Cladera et al. [2 2]). Tanaka et al. [ 23] presented an iron -based polycrystalline SMA (an alloy made of Fe -Ni-Co- Al-Ta-B) that showed extremely high tensile strength (more than 1000 MPa) and a supe relastic strain above 13% at room temperature. In addition, Omori et al. [ 24] discovered another iron -based SMA (Fe -Mn-Al-Ni), with a recoverable strain of more than 5%, showed acceptable superelastic characteristics at room temperature. Nazari and Ghiami Azad et al. [25] numerically studied endplate connections equipped with Fe -based SMA introduced by Omori et al. [ 24]. These connections' moment capacity, stiffness, recentering ability, and energy dissipation were compared with conventional endplate connec tions. Numerical Model Verification The Abaqus finite element software was used to create and analyze 3D models to evaluate the response of extended endplate connections equipped with SMA bolts. Before beginning the in -depth analysis, it is required to f irst validate the SMA connection model. For this purpose, one of the experimentally tested endplate connections is equipped with SMA bolts (Fang et al. [ 26]) is chosen to make a comparison between finite element model and test resu lts. The fourth specimen ("SMA -D10 - 290") is an edge beam -column connection in which a cantilever beam connected to a column by an extended endplate welded to the beam end and HS bolts are substituted by long superelastic SMA bolts. The beam dimensions for the SMA connection specim ens were chosen to guarantee that the beams would stay elastic while applying the cyclic load. Geometric details and mechanical properties of steel materials used in the experimental specimen are depicted in “Figure 1” and “Table 1” , respectively. Table 1. Mate rial properties of steel Mater ial Modulus of Elasticity (GPa) Yield Strength (MPa) Ultimate Strength (MPa) Fracture Strain (%) Steel (beam) 210.1 287.5 452.9 28.3 Steel (end- plate) 205.6 377.5 531.2 24.9 Steel (column) 193.3 306.1 487.4 26 Bilinear stress -strain relationships were considered for steel materials, and the elastic modulus and Poisson ratio were assumed to be 200 GPa and 0.3, respectively. Additionally, combined (isotropic -kinematic) hardening was chosen in the numerical model f or steel material's plasticity and nonlinearity. 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 4 Figure 1.Description of experimental model [ 26] Auricchio's mode l (Auricchio et al. [2 7]) was used to define the superelastic behavior of SMA in the Abaqus software. In Au ricchio's model, the key parameters utilized to characterize a superelastic SMA material are the elastic moduli of SMA in the austenite and martensite phases (EA, EM), transformation stresses (σMs, σMf, σAs, and σAf), maximum transformation strain (ε l), and Poisson ratios (νA, νM). This model indicates that the external force -induced deformations in the SMA material can be entirely recovered after unloading under isothermal circumstances. Furthermore, residual cumulative strains under cyclic loads are omitted, and there is no connection between loading rate and material behavior. In the experimental study, The Ni -Ti bars, similar in type and diameter to the SMA bolts in the tested connections, were subjected to coupon tests at room temperature [2 8]. Based on these results, the SMA nominal properties are mentioned in the “Table 2”. In the numerical model, these engineering values are converted to actual stress -strain values. Table 2. Material properties of SMA bolt used in Abaqus ([26]) Material properties Value Modulus of elasticity EA (GPa) 40 Modulus of elasticity EM (GPa) 40 Austenite -to-martensite starting stress σMs (MPa) 350 Austenite -to-martensite finishing stress σMf (MPa) 450 Martensite -to-austenite finishing stress σAs (MPa) 150 Martensite -to-austenite finishing stress σAf (MPa) 75 Poisson ratios νA, νM 0.33 Transformation strain εL (%) 5 All components were built using solid parts. Large displacements and strains were more precisely considered using nonlinear geometry (an option in Abaqus) throughout loading steps. As mentioned before, corresponding to design details, which beams stayed in the elastic range, buckling did not happen, and regarding [ 29] half modeling of connections considering symmetric boundary conditions can be used without inaccuracy. This method leads to faster computation time without affecting the outcomes. The symmetry plane passes through beam and column center lines. For all contacts, hard contact with a friction coefficient of 0.2 was u sed, which is equivalent to the Class D slip coeffici ent (Eurocode 3 [ 30]). since the weld material is stronger than structural steel, The welded parts were incorporated utilizing surface -based "Tie" constraints. Apart from symmetry, other boundary conditi ons are identical to those in the experimental model. The transitional degrees of freedom were constrained by defining boundary conditions on the column top and bottom surfaces (fixed support condition). Three - dimensional, eight -node linear brick elements with reduced integration were used to mesh each part of the model (C3D8R). SMA bolts were meshed with the same elements but with a nonreduced integration method. Typically, loading consists of two steps. The first step included pre -strain which would resul t in prestresses, and the second step referred to mechanical load (i.e., beam tip drift). By applying a negative temperature gradient to the bolts' shanks, pre-tensioning was accomplished. For SMA material, the thermal expansion coefficient needs to be hig h in order to prevent errors caused by large temperature gradients. Additionally, orthotropic expansion was defined to provide thermal contraction in a single direction (bolt longitudinal axis). The temperature variation was selected so that all the SMA bo lts would experience approximately 230 MPa prestress, about 65% of forward transformation stress. This preload was applied for initial stiffness, self -centering ability, and sufficient friction between the endplate and column face [ 31]. further levels of p retension wou ld result in bolt twisting [2 6]. Speicher et al. [ 30], in another experimental study on the SMA -equipped connections, revealed that pre -strain more than 0.5% would be impractical due to the tendency of bolt shanks to twist. The second step included vertical displacement of the beam tip under the SAC loading protocol [2, 30]. This protocol stipulates several load cycles with increasing inter -story drift levels. The inter -story drift should be calculated by dividing the beam ti p displacement by the distance between the column center and the loading point. It is noteworthy to mention that in the finite element analysis, only one load cycle was conducted for each drift level to increase the computation rate. Some instructions are necessary for finite element analysis to prevent numerical problems. Mesh quality should be evaluated to identify distorting elements (if there are 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 5 any), slave surfaces mesh density should be finer than master surfaces mesh density, each drift level should be applied in several increments according to nonlinearity , etc. Figure 2. Verification of the numerical analysis Figure 3. (a) Deformed shape and von Mises stress contour of the numerical model at 5% drift in comparison with (b) the deformed shap e of the experimental specimen [26] “Figure 2” shows the relationship between the connection plastic rotation for analytical and experimental models and the bending moment at the column face. The plastic rotat ion was calculated by subtracting the beam's elastic rotation from the overall drift. The elastic rotation of the beam is equal to the applied force divided by elastic stiffness derived from a model with rigid bolts ( Fang et al. [26]). Because of the limit ation of the SMA material constitutive model, the hysteresis curve generated by the numerical model does not perfectly match the experiment's outcomes. The differences between the unloading regions of the two curves are primarily due to SMA model idealizations. However, the numerical model accurately predicted the test specimen's overall cyclic response concerning stiffness, strength, energy dissipation, and self - centering ability. “Figure 3” depicts the connection's deformed shape at the ul timate drift level. As shown, the finite element model deformation and the deformation of the experimental specimen closely match. Connection’s General Design Procedure Two beam -column connections were sele cted from NIST benchmarks in order to compare the cyclic response of extended endplate connections equipped with a tested Ni -Ti and an iron -based SMA. Aside from the SMA bolts, AISC design guides were used to design the other parts of the connections. For diminishing prying actions, thick endplates a re chosen in designs. Steel material was A992 (F y= 50 ksi, F u= 65 ksi) and SMA s' properties are listed in “Table 3 ”. Table 3. Material properties of SMAs used in this study Material Ni-Ti [31] Fe-Mn-Al-Ni [24] EA (ksi) 4350 14200 EM (ksi) 2900 8690 σMs (ksi) 44 46 σMf (ksi) 65 58 σAs (ksi) 29 30 σAf (ksi) 15 13 νA, νM 0.33 0.33 εL (%) 5 5 Figure 9. connections' geometry 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 6 Each model's name identification is divided into three parts: connection type, SMA type, and bolt length, respectively. The first part stands for connection type:  EC4 4-bolt  EC4S 4-bolt with endplate stiffeners  EC8S 8-bolt with endplate stiffeners The second part represents the SMA type:  Ni for Ni -Ti and Fe for Fe -based SMAs. The third part introduces SMA bolts length:  L for long bolts (12.5 inches)  I for intermediate bolts (10 inches)  S for short bolts (7.5 inches). These lengths are approximately equal to 50%,40%, and 30% of the beam's depth. There are two connections in which the SMA bolts are one grade smaller (in diameter), in which the letter W follows their names. For instance, the name EC4 -Fe- I represents a 4 -bolt extended endplate connection utilizing Fe -based SMAs in which the shanks are 10 inches. Table 4. beams, columns, and SMA diameters connection beam column bolts' diameters (inch) 4-bolted W24X55 W14X145 1.125 (1 for weaker) 8-bolted W27X114 W18X175 1.25 (1.125 for weaker) SMA Bolts Design Procedure In conventional procedures, plastic hinges develop on the beam ends as the structure experiences seismic excitations. On the other hand, in SMA -equipped connections, the plastic or, more precisely, the superelastic hinge is located at the beam -column interface, where SMA bolt elongation results in a ductile connection. This purpose can be accomplished by choosing lower values for the bolt series' moment capacity relative to the beam flexural resistance. Sultana and Youssef [ 20] in their study on moment -resisting frames equipped with SMA connections, assumed that: 0.8𝑀𝑝𝑏 =2 𝜎𝑚𝑠×𝜋𝑑𝑏2/4×(∑ℎ𝑖) 𝑀𝑝𝑏 is the beam plastic bending moment, 𝑑𝑏 is bolts diameter, 𝜎𝑚𝑠 is forward transformation stress of SMA material, and ℎ𝑖is depicted in “figure 9”. In another study conducted by Roshanfekr Rad et al. [21], the diameter of the bolts was chosen differently: 𝑀𝑦,𝑏𝑒𝑎𝑚 =(𝜎𝑚𝑠+𝜎𝑚𝑓)×𝜋𝑑𝑏2/4×(∑ℎ𝑖) Where 𝑀𝑦,𝑏𝑒𝑎𝑚 is the elastic moment capacity of the beam section and 𝜎𝑚𝑓 is the final stress for forward phase transformation of SMA. This study used the second approach to calculate the Ni -Ti bolts' diameters. Fe -based SMAs with the same diameter as Nitinol bolts were incorporated into the connections to compare Ni -Ti with Fe -based SMA bolts. Bolts diameters are listed in the “Table 3”. All the SMA bolts received a pre -strain of 0.5% at the bolt pre -tensioning stage. SMA shanks were subjected to a negative temperature gradient to make the average longitudinal stress in the center of the bolts equal to 0.5 100𝐸𝐴 (stress at 0.5 % strain). The loading process was carried o ut similar to the SAC loading protocol [ 31] with two additional drift levels (5% and 6%). It should be noted that as a verification model, every drift level was applied for one cycle to reduce computational time. The properties of SMA materials are listed in “Table 3”. Results and Discussions This section presents the cyclic behavior of the six connection models. “Figure 4” illustrates the deformed shape of three models at 6% drift level. Due to limitation of SMA bolts in these type s of connections, designing a connection which deforms without beam yielding at high drift levels isn't practicable. therefore, plastic strains are remained in parts of the beams. “Figures 4-8” illustrate end -plate connections' moment -rotation hysteresis . The sections below discuss about connections' recentering, strength, stiffness and energy dissipation. (a) 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 7 (b) (c) Figure 4. deformed shape of connections at final drift level. (a) EC4-Fe-I, (b) EC4S -Ni-S, and (c) EC8S -Fe-L (a) EC4S -Fe & Ni -I (b) EC4S -Fe & Ni -S (c) EC4 -Fe-I (d) EC8S -Fe & Ni -I (e) EC8S -Fe & Ni -L Figure 5. moment -interstory drift of connections (a) EC4S-Fe & Ni -I (b) EC4S -Fe & Ni -S (c) EC4 -Fe-I (d) EC8S -Fe & Ni -I 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 8 (e) EC8S -Fe & Ni -L Figure 6. moment -rotation joints Figure 7. moment inter-story drifts, EC4S -Fe-I & EC4S -Fe-I-W Figure 8. moment -inter-story drifts, EC8S -Fe-L & EC8S -Fe-L-W recentering An evaluation of the residual drifts can be used to estimate a beam -column connection's self -centering capacity. Following is a definition of the self - centering index ( SCI) of the connections at each inter-story drift level: 𝑆𝐶𝐼𝜃=100 (1−𝜃𝑟 𝜃) Where 𝑆𝐶𝐼𝜃 is the percentage of the connection recentering (self -centering) ability after an inter - story drift level of 𝜃 and 𝜃𝑟 is the residual drift. “Figures 9 a and b ” depict the 𝑆𝐶𝐼𝜃 for 4 -bolted and 8-bolted end -plate connections, respectively. At drifts of 0.02 rad and below, the recentering indices for all connections were more than 96%, meaning that the connections were able to totally restore their initial states after unloading. Although thick end -plate was chosen for EC4 - Fe-I, the recenter ing capacity of this connection isn't satisfactory. by stiffening end -plate the 𝑆𝐶𝐼𝜃 for EC4S -Fe-I significantly improved. as drift levels go further, the recentering ability of connections diminishes due to two main reasons. the first reason is that a s joint exposed to higher levels of rotation, yielding in end -plates stiffeners , which is produced by concentrated compressive forces, increases. also, high level of drifts results in end-plate flexural yielding around farthest bolt rows. the second reason is that when SMA bolts' stresses goes beyond the σmf, beam reaches it is plastic moment capacity and plastic hinge develops on the beam end. Models with weaker SMAs didn't exhibit a considerable difference in 𝑆𝐶𝐼𝜃 due to joint plastic deformation ( i.e., the second reason mentioned above). Overall, intermediate and long bolt lengths for 4 - bolted and 8 -bolted connections seems to be more efficient to reduce residual drifts, respectively. Furthermore, almost all the connections equipped with iron -based S MAs, demonstrated good recentering which weren't far from Ni -Ti connections. (a) (b) Figure 9. 𝑆𝐶𝐼 𝜃 (a) 4-bolted and (b) 8 -bolted strength and stiffness All the connections reached the plastic moment capacity of connected beams at 6% drift. So, all of them are full -strength connections . “Table 4” shows the stiffness of the connections. Longer bolts made connections less stiff, without changing the final moment capacity. if a connection's initial stiffness (K con) for moment - resisting frames is gr eater than 25EI/L (where EI/L is the stiffness of the connected beam and is twice the length of cantilever beam), the connection is in rigid category. on the other hand, when Kcon is less than 0.5EI/L, the connection is regarded to be nominally pinned, and connections with initial 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 9 stiffness between the two limits are considered semi-rigid. Therefore, according to numerical results, these connections are characterized as semi -rigid connections. Table 4. connections' rotational stiffness CONNECTION Kcon (*10^6 kip.in/rad) Kcon/(EI/L) EC4-Fe-I 1.9 17.5 EC4S -Fe-I 2.3 21.2 EC4S -Ni-I 1 9.2 EC4S -Fe-S 2.4 22.2 EC4S -Ni-S 1.54 14.2 EC4S -Fe-I-W 1.9 17.5 EC8S -Fe-I 6.6 20.1 EC8S -Ni-I 3.7 11.3 EC8S -Fe-L 6.4 19.5 EC8S -Ni-L 3.46 10.5 EC8S -Fe-L-W 6.15 18.7 Table 5. dissipated energy dissipated energy (kip.in) CONNECTION total joint EC4-Fe-I 1222 710 EC4S -Fe-I 1394 534 EC4S -Ni-I 1041 522 EC4S -Fe-S 1670 444 EC4S -Ni-S 1370 446 EC4S -Fe-I-W 1197 570 EC8S -Fe-I 4952 1437 EC8S -Ni-I 4064 1341 EC8S -Fe-L 4275 1628 EC8S -Ni-L 3329 1461 EC8S -Fe-L-W 4162 1915 energy absorption The cumulative energy dissipation (the area encompassed by the hysteresis loops) was determined for both joints and total rotations to gain a better understanding of the cyclic performance of the models. “Table 5” illustrates dissipated energy in all models. As expected, joints with shorter bolts showed lower damping . However, overall damping of these models were more than the others with longer SMA bolts. In other words, Beams' plastic hinges increased overall dis sipated energy. EC4-Fe-I dissipated more energy in it's joint in relation to EC4S -Fe-I. this was happened because of plastic deformations occurred in the unstiffened end - plate. Total energy absorption in the second connection was about 14% more than the fi rst one which indicates higher plasticity in the stiffened connection's beam. In all cases, iron -based SMA joints exhibited approximately similar levels of energy dissipation. However, overall energy absorption in Ni -Ti connections was lower numbers. That is induced by significantly higher stiffness of Fe -based SMA bolts. In addition, models incorporating weaker bolts can be used with respect to their capacity to act as full- strength connections (which results in decrease of SMA material usage). In fact, th e final moment capacity of connections for weaker models as others is dependent to beams' plastic moment. Conclusion in this article, the cyclic response of two different beam -to-column connections were numerically investigated to compare the performance of end - plate connections equipped with two different types of superelastic SMA bolts. following conclusions were extracted from the finite element results: - First, comparisons with the test data were used to validate the FE modelling method. It was shown th at the test results and the FE results had acceptable agreement. - since in SMA equipped connections, the ductility and recentering capability are mostly provided by elongation of SMA bolts, incorporating longer SMAs would result in higher superelastic ducti lity and recentering. - Hysteretic curves showed that using SMAs with one grade smaller diameters didn't change the maximum moment capacity of the connections at high drift levels. However, before drifts of about 4%, the resistance of models was significantl y depended on SMAs' diameters. - Iron-based SMA can be a substitute for Ni - Ti. Results showed that the recentering ability of connections equipped with Fe - based SMA bolts, can accommodate self- centering almost equal to connections used Ni-Ti bolts. Also due to higher module of elasticity in iron -based SMAs, connections utilized iron -based bolts had higher levels of stiffness. - Due to high cost of Ni -Ti materials, using iron-based SMAs, which their costs are approximately 0.1 of Nitinols, can have a promising prospect in the future. 2nd International Conference on Industrial Application of Advanced Materials and Manufacturing July. 20th to 21th, Iran University of Science and Technology, Tehran, Iran 10