Experimental investigation of seismic behaviour of low dissipative CFS strap-braced stud walls

Journal of Constructional Steel Research 127 (2016) 92–107

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Journal of Constructional Steel Research

Experimental investigation of seismic behaviour of low dissipative CFS strap-braced stud walls Luigi Fiorino, Maria Teresa Terracciano, Raffaele Landolfo ⁎ Department of Structures for Engineering and Architecture, University of Naples “Federico II”, Naples, Italy

a r t i c l e

i n f o

Article history: Received 25 April 2016 Received in revised form 19 July 2016 Accepted 25 July 2016 Available online 1 August 2016 Keywords: Low dissipative Cold-formed steel Wall tests Strap-braced stud walls Seismic behaviour

a b s t r a c t The adoption of a constructive system in seismic area is related to the clarity and direct applicability of the reference technical code. The use of Cold-Formed Steel (CFS) products for structural applications in seismic area has not been properly contextualized in the current European technical standard. In order to bridge this gap, many national and international companies are promoting experimental activities with the main aim of investigating and interpreting the seismic behaviour of these systems. In this framework, an important collaboration between the University of Naples “Federico II” and the Lamieredil S.p.A. company started in the last years. In particular, the research activity involves the experimental characterization of the seismic response of strap-braced CFS walls by means of monotonic and cyclic tests on full-scale prototypes. Furthermore, in order to evaluate the influence of the main structural components on the global response, the experimental campaign is completed with tests on materials and main connection systems. © 2016 Elsevier Ltd. All rights reserved.

1. Introduction The use of cold-formed steel (CFS) profiles in low-rise residential buildings has increased in European construction sector. The industry interest is related to potentialities offered by this constructive system, which is a competitive solution that meets high structural, technological and environmental performance. In particular, the CFS structures are able to ensure a good structural response in seismic areas, mainly thanks to the lightness that characterizes them. In these structures, the lateral load bearing systems are CFS stud walls, that are generally realized with frames made of CFS profiles braced by sheathing panels or steel straps generally installed in a X configuration. In this last case, the design is usually carried out by neglecting the effect of sheathing panels a considering only the steel elements (so called “all-steel” approach). The unconventionality of these systems has motivated, in recent times, many research groups to carry out several experimental programs, aimed to investigate the seismic performance of CFS strapbraced stud walls (Adham et al. [1]; Serrette [2,3]; Serrette and Ogunfunmi [4]; Fulop and Dubina [5]; Tian et al. [6]; Al-Kharat and Rogers [7]; Casafont et al. [8–10]; Moghimi and Ronagh [11]; Velchev et al.

⁎ Corresponding author at: Department of Structures for Engineering and Architecture, University of Naples “Federico II”, Via Forno Vecchio 36, 80134 Naples, Italy. E-mail addresses: lfi[email protected] (L. Fiorino), [email protected] (M.T. Terracciano), [email protected] (R. Landolfo).

http://dx.doi.org/10.1016/j.jcsr.2016.07.027 0143-974X/© 2016 Elsevier Ltd. All rights reserved.

[12]; Iuorio et al. [13]; Macillo et al. [14]). In particular, the studies were focused on the monotonic and cyclic response of these systems in order to evaluate the seismic response of strap-braced stud walls in terms of strength, stiffness, deformation capacity and energy dissipation. The contribution of the frame without bracing was analyzed by many studies [4,6,7]. Specifically, Tian et al. [6] estimated that a frame, with aspect ratio (height-to-length, 2450 mm × 1250 mm) 2:1, without any bracing system has a lateral strength b 5% of the braced one. Different research included wall specimens made of strap braces on one or both wall sides, different strap dimensions and steel material properties [1,3,4,6,10]. The experimental results highlighted the design of the tension diagonal is a key issue for the seismic response of CFS strap-braced stud walls. The walls braced with steel flat straps installed in an X configuration on both sides showed a better performance than one-side X-braced walls [1,4,6], because the walls braced on one-side only failed by excessive lateral deflection [4] and then the maximum load was reduced by about than 50% respect to walls braced on both sides [6]. Another key issue is the frame-to-strap connection, which highly influences the wall strength and ductility. This issue was investigated by Al-Kharat and Rogers [7], Velchev et al. [12], Casafont et al. [8–10], Moghimi and Ronagh [11] and Iuorio et al. [13]. In particular, Casafont et al. [10] highlighted that the walls should be designed in order to fail for effect of brace yielding followed by strap net-section failure, which is a preferable collapse mode that allows a good wall seismic performance. Furthermore, Velchev et al. [12] showed that the welded and screwed frame-to-strap connections exhibit similar inelastic behaviour

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107 Table 1 Main parameters for the definition of the seismic action. Geographical position

ag [g]

Fo

T*C [s]

SS

ST

Naples

0.17

2.38

0.34

1.20

1.00

ag: peak ground acceleration on bed rock; Fo: spectrum amplification factor; T*C: starting period of the constant speed branch of the horizontal spectrum; Ss: stratigraphic amplification factor; ST: topographic amplification factor.

than those experimentally obtained (q = 2.0 ÷ 2.2). For dissipative walls, the value provided by AISI S213 in case of “Limited ductility” braced walls (q = 2.5) represented a lower limit of the test-based obtained behaviour factor (q = 2.5 ÷ 8.2) [13]. Since guidelines for the seismic design of CFS structures are not provided by the European codes (EN 1998-1 [20]), many national and international companies are promoting experimental activities with the main aim of investigating and interpreting the seismic behaviour of these systems. Part of these studies focuses the attention on the assessment of the seismic behaviour of such construction systems designed according the “all-steel” approach [21]. In this framework, an important cooperation between the University of Naples “Federico II” and Lamieredil S.p.A. company is started in the last years. The main goal of this collaboration is the experimental characterization of global and local seismic response of low dissipative CFS strap-braced walls. This paper presents the experimental campaign and discusses the main results.

2. Seismic design of all-steel CSF systems according to current seismic codes

3m

3rd level

9m 3m

The applicability and the diffusion of a structural system in a seismic area are related to the clarity and the interpretation of technical prescriptions. In general, seismic codes classify buildings on the base of the ductility requirements and the dissipation capacity of a given seismic resistant system. The behaviour factor q (according to the European terminology) or the seismic response modification factor R (according to the USA terminology) is the main design parameter that quantifies the inelastic capacity of the structural system and it represents a fundamental issue to deepen when design prescriptions for a new seismic resistant system are going to be proposed. The most advanced code covering the seismic design of diagonal strap braced (i.e, all-steel) CFS systems are the “North American Standard for Cold Formed Steel Framing - Lateral Design” (AISI S213 [17]). The AISI S213 is developed by the American Iron and Steel Institute Committee on Framing Standards and it codifies the design under wind and seismic loads of different lateral resistant CFS systems for Canada, Mexico and United States. Both sheathed shear walls and strap-braced systems are considered in the

2nd level

3m

12.2 m

if properly designed and detailed to avoid strap net section fracture before the brace yielding. The methodologies used for designing CFS strap-braced walls are widely described in Kim et al. [15], Al-Kharat and Rogers [7], Velchev et al. [12] and Macillo et al. [14]. Kim et al. [15] reported the experimental results of a full-scale two storey one-bay structure tested on shake table. The design of the specimen followed the TI 809-07 [16] by assuming a seismic response modification factor (R) equal to 4, which corresponds to the behaviour factor (q) in the European code terminology. Non dissipative elements (connections, studs and anchors to the lower beam and to upper floor slabs) were designed according to TI 809-07 [16], considering the maximum over-strength of the straps, in such a way that the specimen behaves in the desired ductile manner. Test results showed that cold-formed steel structure behaved very well under seismic loading. The experimental campaign carried out by Al-Kharat and Rogers [7] included both elastic and dissipative walls. For the elastic approach, the behaviour factor was set equal to 1.6 according to the prescription of AISI S213 [17] for “Conventional Construction”. Furthermore, dissipative walls were detailed following a proposed seismic capacity design approach similar to CSA S16 (2005) [18] and by assuming a R factor equal to 4.0 according to ASCE7-05 [19]. The experimental performance of the tested strap-braced walls was affected by the constructive detail of the hold-down devices, which in some cases did not allow reaching or maintaining of the strap yield capacity and thus involved a severely overall system ductility reduction. “Test-based” behaviour factor values of 3.65, 2.11 and 1.72, obtained by authors, indicated low ductility levels, which were not compatible with the R factor value (4.0) provided by ASCE7-05. Velchev et al. [12] performed monotonic and cyclic tests on full-scale wall specimens, designed according to the capacity design requirements of AISI S213 [17] and considering a R factor equal to 2.5. In particular, three factored lateral load levels were used for designing the tested walls: 20 kN (light), 40 kN (medium) and 75 kN (heavy). On the basis of test results, R factors were evaluated and the obtained values (minimum and maximum values equal to 1.6 and 10.4, respectively) were in most cases greater than those used in design phase. Finally, in Macillo et.al [14] three lateral resisting systems made of CFS strap-braced stud walls were designed according to elastic or dissipative design approaches and were tested [13]. The wall designed according to an elastic approach (q = 1) was representative of a onestorey building located in a medium-low seismicity zone. The other two walls were designed according to a dissipative approach, by considering the behaviour factor given by AISI S213 for “Limited ductility” in Canada (q = 2.5) and by using the capacity design rules based on Eurocode approach for traditional steel braced structures, whenever applicable. In the case of elastic light walls, the behaviour factor value adopted in design phase (q = 1) and the one proposed by AISI S213 for “Conventional Construction” category (q = 1.6) were always smaller

93

18.1 m CFS Strap-Braced Stud Wall

Fig. 1. Schematic view of the case study building.

1st level

94

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107 Table 3 Test matrix. Structure

WALLS

Label Representative of No. monotonic tests No. cyclic tests

WHE First level wall (W1) 2 3

Material STEEL Label S280–1.5 (steel grade - thickness in mm) Representative of WLE CFS profiels

Fig. 2. Elastic acceleration spectrum adopted for the case study building.

standard. In particular, with regard to the all-steel solution, the standard contains special requirements for seismic design, such as the values of the seismic response modification factor (R), aspect ratio limitations, capacity design rules for non-dissipative elements. For the definition of the force reduction factor (R) the AISI defines two categories of seismic-resistant systems for diagonal strapbraced wall. For the first category, special seismic requirements (capacity design rules) are not required and the seismic resistant system is not specifically detailed for ductile performance. In this case, in the USA and Mexico, the force reduction factor should be taken equal to or b3, while in Canada, it should be taken equal to or b1.6. For the second category, the rules of capacity based design approach apply. Therefore, for this case, in the USA and Mexico, the force reduction factor can be taken N3, in accordance with the applicable building code, e.g. the American code “Minimum Design Loads for Buildings and Other Structures” (ASCE/SEI 7–10 [22]) provides a factor value equal to 4 for diagonal strap braced walls. In Canada, for the all-steel systems designed according to capacity design, an R factor equal to 2.5 can be used. Nowadays in Europe there are not specific prescriptions for the seismic design of CFS structures. Indeed, for seismic resistant steel buildings the EN 1998-1 [20] defines three structural ductility classes: high (DCH), medium (DCM) and low (DCL). Structural systems belonging to DCH and DCM classes have a higher ability to dissipate energy and are designed to resist seismic actions taking into account their inelastic capacity. The DCL class structures have a low dissipative behaviour and their design is carried out without taking into account significant non-linear behaviour. In this case, the recommended value for the behaviour factor (q) is ≤ 1.5; the resistance of members and connections should be evaluated in accordance with EN 1993-1-1 [23] without any additional requirement; and no restrictions on crosssection class are recommended (also members with slender crosssection can be used). Therefore, according to the current provisions of EN 1998, the seismic design of all-steel CFS walls is possible by considering them as common steel structures made of Class 4

WLE Third level wall (W3) – 1

S280–2

S275–20

S235–8

WHE CFS profiels

Bottom plate of HD-H hold-down 3

U-profile of HD-H hold-down

No. tests

3

3

3

Component Typology Representative of No. tests

SCREWS hexagonal head self-drilling screws steel-to-steel wall connection 3

Component Label Representative of No. tests

JOINTS between GUSSETS PLATE and STRAP-BRACE CHE CLE WHE wall WLE wall 3 3

Component Label Representative of No. tests

HOLD-DOWN HD-H WHE wall 2

HD-L WLE wall 2

cross-sections belonging to the DCL with a behaviour factor ≤ 1.5. The prescriptions of the Italian construction code [24] are similar to those provided by EN 1998, without specific rules for the seismic design of CFS structures. In the case of elastic approach (behaviour factor q or seismic response modification factor R assumed equal to 1), the wall lateral resistance can be evaluated as the smaller resistance associated to all possible collapse mechanisms of the elements that compose it (studs, tracks, diagonals strap-braced, strap connections, tension and shear anchors) without any expedient which promotes a ductile collapse mechanism. Therefore, the lateral strength of the wall (Hc) can be computed according to following equation [14]: H c ¼ min H c;d ; Hc;c ; H c;s ; H c;t ; H c;a




ð1Þ

where, Hc,d is the wall strength associated to the yielding of diagonal steel straps; Hc,c is the wall strength associated to the diagonal-to-strap connections failure; Hc,s is the wall strength associated to the studs buckling; Hc,t is the walls strength associated to the tracks buckling; and Hc,a is the wall strength associated to the tension or shear anchors failure. 3. Case study

Table 2 Design hypotheses and results for selected wall configurations. Wall configuration

W1

W2

Storey Seismic action on single wall (Hd) [kN] Design lateral wall resistance (Hc) [kN] Lateral wall stiffness (k) [kN/mm] Predicteda collapse mechanism

First 121 152 6.8 Local buckling of tracks

Second Third 97 54 98 73 5.5 4.2 Diagonal net area failure

W3

a Evaluated according to the EN 1993-1-3 [23] through the methodology illustrated in [14] by using the design mechanical properties of the materials.

For planning the experimental campaign and defining the configurations of diagonal strap-braced walls to be examined, a residential building was considered as case study. The studied building, with rectangular plan, covered an area of 220 m2 and had three storeys with a storey height of 3.00 m. From seismic point of view, the building was regular both in plan and elevation. In fact, the seismic resistant systems, which corresponded to ten CFS strap-braced stud walls per each direction, had a plan symmetrical distribution (Fig. 1). The building was designed considering environmental loads acting in Naples, South Italy, which is characterized by medium-low intensities of snow and seismic loads. Steel sheet – lightweight concrete composite floors and roof were

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107

Fig. 3. Schematic drawings of the wall configurations: a) elastic heavy wall (WHE); b) elastic light wall (WLE); c) blocking members detail; d) track reinforcement detail.

95

96

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107

Table 4 Nominal design dimensions and material properties of the tested walls. WHE (W1a)

WLE (W3b)

Section [mm] Studs Track Diagonal Straps Gusset plates Track reinforcements Blocking

Section [mm] c

C150 × 60 × 20 × 1.5 U153 × 60 × 1.5d 160 × 1.5e 300 × 300 × 1.5f C150 × 60 × 20 × 1.5c C150 × 60 × 20 × 1.5c U153 × 60 × 1.5d 60 × 1.5e

Flat strap Hold-down

C150 × 60 × 20 × 2 U154 × 60 × 2d 240 × 2e 385 × 385 × 2f C150 × 60 × 20 × 2c C150 × 60 × 20 × 2c U154 × 60 × 2d 60 × 2e Purposely designed welded steel hold-down

Screws Shear anchors Hold-down to chord stud fasteners Hold-down to steel beam fasteners

6.3 × 40g hexagonal head self-drilling screws M10 class 8.8 bolts spaced at 200 mm on centre No.4 M18 bolts No.4 M14 bolts No. 1 M24 bolt No. 1 M20 bolt

a b c d e f g

Steel Grade c

S280GD S280GD S280GD S280GD S280GD S280GD S280GD S235JR S275JR

Class 8.8 Class 8.8

WHE specimens were representative of the walls W1 located at the first level of the case study building. WLE specimen was representative of the walls W3 located at the third level of the case study building. C-section: outside-to-outside web depth x outside-to-outside flange size x outside-to-outside lip size x thickness. U-section: outside-to-outside web depth x outside-to-outside flange size x thickness; Width × thickness. Height × width × thickness. Diameter × length.

assumed for the evaluation of floor dead loads, equal to 1.02 kN/m2. Live loads for residential buildings equal to 2.00 kN/m2 were considered for both floors and roofs. The snow load was calculated for the assumed geographic location according to Italian construction technical code [24] and it was equal to 0.48 kN/m2. The seismic action and the design spectra were defined according to the Italian construction technical code, which provides the reference peak ground acceleration on bed rock on the basis of geographical position of the construction site, which is equal to 0.17 g (for Naples). The assumed foundation soil was type B, according to EN 1998-1 [20], which corresponds to loose to slightly cemented pyroclastic deposit. The main parameters for the calculation of the seismic action for earthquakes having 10% probability of exceedance in 50 years, are

summarized in Table 1, whereas the assumed elastic acceleration spectrum is shown in Fig. 2. The seismic design was carried out through a linear dynamic analysis and considering a non-dissipative (elastic) behaviour (q = 1). In the analysis, the floors were assumed as rigid diaphragms and the effects of accidental eccentricity of the mass were neglected. All diagonal strapbraced walls had dimension 2400 mm × 2700 mm (length × high). As it is well known, the lateral seismic force demand decries form first (ground) to third level, therefore three different walls (W1, W2, and W3) were designed according to the procedure described in [14] for the elastic approach. Table 2 shows the results of the walls design in terms of wall components, seismic force demand, seismic strength capacity and lateral

Fig. 4. Wall specimen WLE - general view and corner detail.

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107

97

Fig. 5. Wall specimens WHE - general view and corner detail.

stiffness. The walls were designed without following any prescription aimed at avoiding brittle failure mechanisms, with the only exception of the brittle failure of the fasteners, for which the prescription given by EN 1993-1-3 [23] to provide an adequate overstrength for the shear fastener rupture was applied. As a consequence, the collapse mechanism predicted in the design phase was the failure of the diagonal net area at the fastener holes location for the walls W2 and W3 and the local buckling of the tracks for the wall W1.

4. Experimental campaign The lateral response of these systems was investigated through the execution of two monotonic and four cyclic tests on full-scale wall specimens, representative of the first and third level of the case study building. Moreover, taking into account that materials and components influence the wall global seismic response, the local response of the walls was investigated by means of twelve tension tests on steel

Fig. 6. Wall tests - test set-up and instrumentations.

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L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107

0.0

Table 5 Results of monotonic wall tests. Label

Hp [kN]

dp [mm]

Hy [kN]

dy [mm]

Δm [mm]

ke [kN/mm]

ke,o [kN/mm]

WHE-M1a WHE-M2a EXP(AV)b THEc EXP(AV)/THE

187.1 185.7 186.4 186,0 1.0

52.9 58.7 55.8 – –

174.8 164.1 169.4 – –

42.1 38.9 40.5 – –

66.7 67.6 67.2 – –

5.6 4.6 5.1 6.8 0.75

9.3 8.0 8.7 6.8 1.27

1.0

H [kN]

200 Hy

3.0

4.0

d/h [%]

Hp

150 100

ke,THE ke,o

50

ke d [mm]

0 dy

0 materials, three shear tests on screws, six shear tests on screwed connections between gusset plates and strap-bracings and four tension tests on hold-down devices. The experimental campaign is summarized in Table 3. All tests were carried out in the laboratory of the Department of Structures for Engineering and Architecture of the University of Naples Federico II.

5. Tests on walls 5.1. Wall specimens In order to investigate the lateral response of the selected strapbraced wall configurations, six tests on full-scale 2400 mm long and 2700 mm high wall specimens were carried out. In particular, two monotonic and three cyclic tests were carried out on five nominally identical specimens, named WHE (Fig. 3a), which were representative of the walls located at the ground floor of the studied building (W1); whereas one cyclic test was carried out on a specimen, named WLE (Fig. 3b), which was representative of the walls located at the third level (W3). The wall framing was made with stud members, having lipped channel sections (C-sections), spaced at 600 mm on the centre and connected at the ends to track members, having unlipped channel sections (U-sections). Chord studs were composed by double C-sections coupled back-to-back. In order to reduce the unbraced length of the chord and interior studs, flat straps were placed at the mid-height of the wall specimens and were screwed to blocking members (Fig. 3c) placed at the ends of walls. The tracks were reinforced by C-section profiles assembled in a box section (Fig. 3d). Hold-down devices were used to anchor chord stud ends and were placed within the lower and upper tracks at the four corners of the walls. In particular, the hold-down devices were connected to the studs by four M14 or M18 class 8.8 bolts and to the beams of the testing frame by one M24 or M20 class 8.8 bolt, for WLE or WHE, respectively. The upper and bottom tracks were

5.0

Hp,THE

a

WHE specimens were representative of the walls W1 located at the first level of the case study building. b EXP(AV): average experimental values. c THE: theoretical values.

2.0

50dp

Δm

100

150

Subscript THE stands for theoretical prediction Fig. 8. Monotonic test on WHE-M1 specimen - load vs. displacement curve.

connected to the both loading (top) and bottom beams of the testing frame by M10 class 8.8 bolts spaced at 200 mm on the centre. The wall specimens were completed with strap braces installed in X configuration on both sides and connected to the wall framing by gusset plates. All the steel members were made with S280 GD steel grade (characteristic yield stress fy = 280 Mpa and characteristic ultimate stress fu = 360 Mpa). All screw fasteners were 6.3 × 40 mm (diameter × length) hexagonal head self-drilling screws. Table 4 lists the nominal design dimensions and material properties of the tested walls. Some photos of the specimens with the corresponding corner details are provided in Fig. 4 and Fig. 5.

5.2. Test set-up, instrumentation and loading protocol Tests on full-scale wall specimens were carried out by using a specifically designed testing frame for in-plane horizontal loading. Since these kind of structures are not affected by P-D effects [25] no gravity loads were applied on the wall prototypes. Horizontal loads were transmitted to the upper wall track by means of a steel loading beam with a 200 × 120 × 10 mm (width × height × thickness) rectangular hollow section. The wall prototype was restrained to the laboratory strong floor by the bottom beam having a 300 × 180 × 30 (width × height × thickness) built-up rectangular hollow section of testing frame. The

0.0 200 Hy

1.0

2.0

3.0

4.0

Hp

H [kN]

5.0 d/h [%] Hp,THE

150 100

ke,THE ke,o

50

ke d [mm]

0 0

dy

50 dp

Δm

100

150

Subscript THE stands for theoretical prediction Fig. 7. Wall tests - cyclic protocol.

Fig. 9. Monotonic test on WHE-M2 specimen - load vs. displacement curve.

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107 Table 6 Phenomena observed during monotonic tests for different inter-storey drift levels vs. expected collapse mechanisms. Phenomenon observed during the test

-8.9 -7.8 -6.7 -5.6 -4.4 -3.3 -2.2 -1.1 0.0 1.1 2.2 3.3 4.4 5.6 6.7 7.8 8.9 120 H [kN] d/h [%] 80

Inter-storey drift levels WHE-M1 (W1a)

Local buckling of the tracks Squashing of the stud ends Out-of-plane deformation of the gusset plate Gusset-to-track connection failure Expectedb collapse mechanism

99

WHE-M2 (W1a)

1.5% 1.3% 1.6% 1.4% 2.0% 2.0% 2.3% 2.4% Local buckling of tracks

40 d [mm]

0 -240

-200

-160

-120

-80

-40

0

40

80

120

160

200

240

-40

a

WHE specimens were representative of the walls W1 located at the first level of the case study building. b Evaluated according to the EN 1993-1-3 [23] through the methodology illustrated in [14] by using the average experimental mechanical properties of the materials.

-80

-120

out-of-plane displacements of the wall were avoided by two lateral supports realized with vertical HEB 140 profiles. The tests were performed by using a hydraulic actuator having a stroke displacement of 500 mm and a load capacity of 500 kN. A sliding-hinge was placed between the actuator and the tested wall in order to avoid the transmission of vertical load components between the actuator and the specimen. Wall displacements were measured by five LVDTs, as shown in Fig. 6. In particular, three LVDTs (P1, P2 e P3) were used to measure the horizontal displacements and two LVDTs (P4, P5) were used to measure the vertical displacements. In particular, the vertical LVDTs were positioned for evaluating the axial deformations of the hold-down devices. The recorded deformations are very small with values ranging between 1 mm and 20 mm. The strain of the diagonal straps was recorded by means of three

Fig. 11. Cyclic test on WLE-C1 specimen - load vs. displacement curve.

strain-gauges, placed at the centre (SG2) and at the two ends of the strap (SG1 and SG3). Monotonic tests were carried out with displacements imposed at a rate of 0.10 mm/s until the collapse of specimens occurred. The data were recorded with a sampling frequency equal to 10 Hz. Cyclic tests were carried out by adopting a loading protocol known as “CUREE ordinary ground motions reversed cyclic load protocol” developed for wood walls by Krawinkler et al. [26]. The cyclic loading test protocol consists of a series of stepwise increasing deformation cycles. The displacement amplitudes of each cycle are defined starting from a reference

Fig. 10. Phenomena observed during monotonic wall tests: a) local buckling of the tracks; b) squashing of the stud ends; c) out-of-plane deformation of the gusset plate; d) gusset-to-track connection failure.

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-8.9 -7.8 -6.7 -5.6 -4.4 -3.3 -2.2 -1.1 0.0 1.1 2.2 3.3 4.4 5.6 6.7 7.8 8.9 240 H [kN] d/h [%] 200 160

160

120

120

80

80 40

40

d [mm]

0 -240

-200

-160

-120

-80

-8.9 -7.8 -6.7 -5.6 -4.4 -3.3 -2.2 -1.1 0.0 1.1 2.2 3.3 4.4 5.6 6.7 7.8 8.9 240 H [kN] d/h [%] 200

-40 0 -40

40

80

120

160

200

240

-240 -200 -160 -120

-80

0 -40 0 -40

-80

-80

-120

-120

-160

-160

-200

-200

-240

-240

d [mm] 40

80

120

160

200

240

Fig. 12. Cyclic test on WHE-C1 specimen – load vs. displacement curve.

Fig. 14. Cyclic test on WHE-C3 specimen - load vs. displacement curve.

deformation Δ = γ Δm, where the value of Δm = 67.2 mm is calculated on the basis of monotonic test results carried out WHE specimens (Table 5), as the displacement corresponding to a load equal to 80% of the maximum load on the post-peak branch of the response curve (conventional ultimate displacement) and γ is assumed equal to 0.6. The γ value is adopted for considering the difference between the deformation capacities obtained by monotonic and cyclic tests in which the cumulative damage lead to earlier strength deterioration. For tested walls the cyclic protocol involved displacements at a rate of 1.0 mm/s, for displacements up to 6.1 mm, 2.0 mm/s for displacements in the range from 6.1 to 30.2 mm, and 4.0 mm/s for displacement larger than 30.2 mm. The cyclic protocol adopted for all tests is shown in Fig. 7.

• Δm: ultimate displacement corresponding to a load equal to 0.80∙Hp on the post-peak branch of the response curve; • ke: conventional secant elastic stiffness defined as the secant stiffness at 40% of the maximum strength; • ke,o: initial tangent elastic stiffness corresponding to the tangent to initial part of the response curve

5.3. Monotonic tests Figs. 8 and 9 show the experimental responses for the specimens WHE-M1 and WHE-M2, in terms acting load (H) vs. top wall displacement (d) curves, in which the top wall displacement was measured by the LVDT P1. The parameters used to describe the experimental behaviour are the following: • Hp: wall strength corresponding to the maximum recorded load; • dp: displacement corresponding to Hp; • Hy: wall conventional yield strength defined at the knee of linear branch of response curve; • dy: displacement corresponding to Hy;

-8.9 -7.8 -6.7 -5.6 -4.4 -3.3 -2.2 -1.1 0.0 1.1 2.2 3.3 4.4 5.6 6.7 7.8 8.9 240 d/h [%] H [kN] 200 160 120 80 40 -240

-200

-160

-120

-80

0 -40 0 -40

d [mm] 40

80

120

160

200

-80 -120 -160 -200 -240 Fig. 13. Cyclic test on WHE-C2 specimen - load vs. displacement curve.

240

Table 5 provides the above defined parameters obtained from the experimental results, together the theoretical expected values of the wall strength and stiffness. The latter were evaluated according to the EN 1993-1-3 [23] through the methodology illustrated in [14] by using the average experimental mechanical properties of the materials. The above defined parameters are also shown in Figs. 8 and 9. Moreover, the inter-storey drift levels, which are defined as the ratio between the top wall displacement (d) and wall height (h) set equal to 2700 mm, are also provided in these figures. Test results showed small differences (3%) in terms of strength between the two nominal identical specimens, whereas larger differences (21%) were obtained for stiffness values. Furthermore, the results highlighted that the theoretical values of stiffness overestimated of 33% the average value of the conventional secant elastic stiffness and they underestimated of 27% the average value of the initial tangent elastic stiffness. According to the expected failure mechanism (Table 6), the observed failure modes for both monotonic tests on WHE configurations were the local buckling of the tracks (Fig. 10a). In particular, the local buckling of the tracks occurred at the elastic branch of the response curve, which corresponded to an inter-storey drift ratio of about 1.4%. Also the squashing of the stud ends was observed at this stage (Fig. 10b). Furthermore, the out-of-plane deformation of the gusset plate (Fig. 10c), related to tension diagonal, was observed at the reaching the wall strength (Hp), which corresponded to an inter-storey drift ratio of 2.0%, whereas the gusset-to-track connection failure (Fig. 10d) occurred in correspondence of a significant reduction of strength on the response curve post-peak branch, which corresponded to an interstorey drift ratio of about 2.3%. 5.4. Cyclic tests Figs. 11–14 show the acting load (H) versus the measured displacement (d) curves for the WLE-C1, WHE-C1, WHE-C2 and WHE-C3 specimens, together positive and negative envelope curves. The results of the cyclic tests are shown in Table 7, in which the experimental parameters were evaluated on positive and negative envelopes and the theoretical (expected) values of the wall strength and stiffness are the same ones provided for the monotonic tests.

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Table 7 Results of cyclic wall tests. Label

HP [kN]

b

WLE-C1 EXP(pull,push,AV)c THEd EXP(pull,push,AV)/THE WHE-C1e WHE-C2e WHE-C3e EXP(AV)f EXP(pull,push,AV)c THEd EXP(pull,push,AV)/THE a b c d e f

dP [mm]

Hy [kN]

Δm [mm]

dy [mm]

ke [kN/mm]

ke,o [KN/MM]

Pusha

Pulla

Pusha

Pulla

Pusha

Pulla

Pusha

Pulla

Pusha

Pulla

Pusha

Pulla

Pusha

Pulla

103.7 103.1 101 1.02 199.7 198.7 204.6 201.0 198.9 186 1.07

102.5

59.6 65.7 – – 52.9 52.5 60.0 55.1 61.2 – –

71.7

78.1 79.9 – – 182.0 191.0 187.2 186.7 181.4 – –

81.8

23.2 29.8 – – 42.9 42.3 39.8 41.7 42.8 – –

36.3

174.5 174.5 – – 70.4 59.3 115.5 81.7 86.4 – –

–

4.1 3.4 4.2 0.81 5.1 7.0 6.2 6.1 5.6 6.8 0.82

2.7

4.2 4.6 4.2 1.10 8.4 8.6 7.9 8.3 7.9 6.8 1.16

5.0

201.1 199.3 190.2 196.9

70.8 72.8 58.2 67.3

191.4 190.2 146.7 176.1

49.1 50.3 32.1 43.8

81.3 100.8 – 91.0

4.8 5.2 5.5 5.2

7.9 8.2 6.6 7.6

Push: positive displacements; pull: negative displacements. WLE specimen was representative of the walls W3 located at the third level of the case study building; EXP(AV): average experimental values; THE: theoretical (expected) values. WHE specimens were representative of the walls W1 located at the first level of the case study building. EXP(pull,push,AV): average of results obtained for push and pull loading phases;

Table 8 Phenomena observed during cyclic tests for different inter-storey drift levels vs. expected collapse mechanisms. Phenomenon observed during the test

Local buckling of the tracks Squashing of the stud ends Out-of-plane deformation of the gusset plate Gusset-to-track connection failure Local buckling of the chord studs Combined bending and compression axial load failure of the chord studs a Observed in the pull loading phase only b Observed in the push loading phase only Expectedc collapse mechanism a b c

Inter-storey drift levels WLE-C1 (W3a)

WHE-C1 (W1b)

WHE-C2 (W1b)

WHE-C3 (W1b)

1.2% 1.2% 2.2% – 2.2% 5.0%

1.5% 1.6% 1.6% 2.0% – –

1.6% 1.6% 1.6% 1.9% – –

1.5% 1.6% 2.2% 4.9%a – 4.3%b

Yielding of diagonal straps

Local buckling of tracks

WLE specimen was representative of the walls W3 located at the third level of the case study building; WHE specimens were representative of the walls W1 located at the first level of the case study building; Evaluated according to the EN 1993-1-3 [23] through the methodology illustrated in [14] by using the average experimental mechanical properties of the materials.

The results showed that the strength and stiffness values recorded in the pushing (positive) phase with respect to the pulling (negative) phase had maximum differences of 6% and 35%, respectively, except

for a stiffness variation of 52% for WLE-C1 specimen. The ratios between the average experimental and theoretical values highlighted that the experimental strength values were always slightly higher than the

Fig. 15. Phenomena observed during cyclic wall tests: a) local buckling of the tracks; b) squashing of the stud ends; c) out-of-plane deformation of the gusset plate; d) local buckling of the chord studs; e) gusset-to-track connection failure; f) combined bending and compression axial load failure.

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Fig. 16. Materials test- stress vs. strain curves.

theoretical predictions, with a maximum difference of 7%. Furthermore, the results in terms of stiffness highlighted that the theoretical values overestimated of 24% the average value of conventional secant elastic stiffness and they underestimated of 16% the average initial tangent elastic stiffness. The failure mode observed for all tests was due to the interaction of different phenomena (Table 8 and Fig. 15). In particular, the tests on WHE-C1 and WHE-C2 walls showed a similar behaviour. The first phenomena were the local buckling of the tracks (Fig. 15a), squashing of the stud ends (Fig. 15b) and out-of-plane deformation of the gusset plate (Fig. 15c), which corresponded to the end of elastic behaviour. The failure of the gusset-to-track connections (Fig. 15d) was observed at the reaching of the maximum wall strength. After the failure of the gusset-to-track connections the strength decreased suddenly. The cyclic test WHE-C3 presented a different experimental response, with a relatively higher deformation capacity compared with the tests WHE-C1 and WHW-C2, especially for the pull loading phase. The first phenomena were the local buckling of the tracks and squashing of the stud ends, which corresponded to the beginning of the inelastic response. These phenomena were followed by the out-of-plane deformation of the gusset plate for larger displacements. In the push loading

phase, the significant decreasing of strength was caused by the global buckling of the chord studs (Fig. 15f), whereas the failure of gusset-totrack connections was observed in the push loading phase at final stage of the test. The experimental responses of the tests WHE-C1 and WHE-C2 showed some differences compared to that of the test WHE-C3. In fact, the specimens WHE-C1 and WHE-C2 showed the failure of the gusset-to-track connections for inter-storey drift ratios of about 2%, whereas for the tests WHE-C3 the combined bending and compression axial load failure of the chord studs was observed only in the push loading phase for an inter-storey drift ratio of about 4%, whereas the failure of the gusset-to-track connections occurred only in the pulling loading phase for larger inter-storey drift ratio, equal to about 5%. Since the gusset-to-track connection collapse and the combined bending and compression axial load failure of the chord studs produced a significant reduction in the load transfer between tension diagonal straps and compressed stud, these events corresponded to a substantial reduction of the wall load bearing capacity. Therefore, the occurrence of these phenomena explains the higher deformation capacity of the wall WHE-C3 compared with the walls WHE-C1 and WHE-C2. For the test WLE-C1 the first phenomena were the local buckling of the tracks and squashing of the stud ends (for an inter-storey drift ratio of 1.2%) followed by the out-of-plane deformation of the gusset plate and local buckling of the chord studs (Fig. 15d) (for an inter-storey drift ratio of 2.2%). For this test the combined bending and compression axial load failure of the chord studs appeared only for very high displacements (for inter-storey drift ratio of 5%), whereas the gusset-totrack connection failure was not observed. As a result, this wall exhibited the best response in terms of deformation capacity, with apparent reductions of the load bearing capacity observed only for inter-storey drift ratios higher than 5%. As a conclusion, the wall inelastic behaviour is affected by non-ductile phenomena, as the gusset-to-track connection failure and the combined bending and compression axial load failure. This condition is typical of structural elements designed without capacity design criteria. Finally, the comparison between monotonic and cyclic test results for the WHE walls revealed that the average experimental strength recorded under monotonic loads is lower than the ones recorded in the cyclic tests (with maximum variations of 8%), whereas the average stiffness for monotonic tests is higher of only 2% than the cyclic one.

Table 9 Results of steel material tests. Nominal values

Experimental results

Stell grade

fy,na [MPa]

fu,nb [MPa]

fu,n/fy,na,b

Thickness [mm]

Label

fy,expc [MPa]

fu,expd [MPa]

fy,exp./fy,na,c

fu,exp./fu,nb,d

fu,exp./fy,expc,d

S280GD

280

360

1.3

1.5

S280–1.5-01 S280–1.5-02 S280–1.5-03 AVe S280–2.0-01 S280–2.0-02 S280–2.0-03 AVe S275–20-01 S275–20-02 S275–20-03 AV(AV)e S235–8.0-01 S235–8.0-02 S235–8.0-03 AVe

283.10 287.28 299.53 289.97 346.29 330.87 353.17 343.45 303.58 341.22 315.31 320.04 253.72 266.41 280.88 267.00

391.76 392.31 416.03 400.03 418.66 409.47 431.14 419.75 465.50 473.01 465.61 468.04 360.83 381.44 391.48 377.92

1.01 1.03 1.07 1.04 1.24 1.18 1.26 1.23 1.10 1.24 1.15 1.16 1.08 1.13 1.20 1.14

1.09 1.09 1.16 1.11 1.16 1.14 1.20 1.17 1.08 1.10 1.08 1.09 1.00 1.06 1.09 1.05

1.38 1.37 1.39 1.38 1.21 1.24 1.22 1.22 1.53 1.39 1.48 1.46 1.42 1.43 1.39 1.42

2.0

S275JR

275

430

1.6

20

S235JR

235

360

1.5

8

a b c d e

fy,n: nominal yield stress. fu,n: nominal ultimate stress. fy,exp.: experimental yield stress. fu,exp.: experimental ultimate stress. AV: average of the experimental values.

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103

Fig. 17. Test set-up for shear tests on screws.

14

between experimental and nominal values of yield and ultimate stresses and ratio between experimental ultimate and yield stresses. The average experimental values of the yield stress were always higher than nominal values (3,5%, 22%, 16%, 14% for S280-1.5, S2802.0, S275-20.0, S235-8.0, respectively). Also the results in terms of average experimental ultimate stress showed always an increment respect to nominal values (11%, 17%, 9%, 5% for S280-1.5, S280-2.0, S275-20.0, S235-8.0, respectively). In addition, the ratios between nominal ultimate and yield stresses were always higher than the ratios between the relevant experimental values (of 7%, 10%, 6% for S280-2.0, S27520.0, S235-8.0, respectively), except for the specimen S280-1.5, in which the experimental ratio was larger of 6% compared to the nominal one.

Sp-6.3-01 Sp-6.3-02 Sp-6.3-03

F [kN]

12 10 8 6 4 2 d [mm]

0 0

1

2

3 6.2. Test on screws

Fig. 18. Screw tests - load vs. displacement curves.

Therefore, for the investigated cases were not observed appreciable degradation effects of the cyclic action for the walls tested with the adopted loading protocol. 6. Tests on material and components 6.1. Test on materials Material coupons of frame members and hold-down devices were subjected to conventional tension tests according to EN ISO 6892-1 [27]. In particular, tests were performed on four specimen types: S280-1.5, S280-2.0, S275-20.0 and S235-8.0. Fig. 16 shows the experimental response curves obtained for all test on materials, whereas Table 9 shows the nominal yield stress (fy,n), nominal ultimate stress (fu,n), experimental yield stress (fy,exp), experimental ultimate stress (fu,exp) and average values of experimental results. In addition, Table 9 shows the ratio between nominal ultimate and yield stresses, ratio

The self-drilling screws adopted for steel-to-steel connections in the investigated configurations of CFS walls were tested in order to evaluate the shear strength. Therefore, three shear tests were carried out on 6.3 mm diameter hex washer head screws. The shear tests were performed by adopting an ad hoc experimental test set-up proposed by Fiorino et al. [28]. In particular, test is performed by passing the screw through three predrilled steel plates set in contact and placed in such a way that the force could be perfectly in line with the screw. Furthermore, the screw drills an additional squared steel plate with thickness 1 mm, in order to avoid the screw pull out during the test. The steel plates are heat treated with hardness values ranging between 50 and 55 HRC, that assure higher strength against plate bearing. Moreover, the plate surfaces have been adequately smoothed in order to reduce the friction and to avoid any loading dissipation. The applied compression load is transferred to the screw through two screw surfaces, therefore the obtained shear strength has been divided between the two tested screw resistant sections (Fig. 17). The parameter used to describe the experimental shear behaviour is the maximum load (Fp) obtained for a single tested screw resistant section and defined on the load (F) vs. displacement (d) curve (Fig. 18).

Table 10 Results of screw tests. Screw typology

Diameter [mm]

hexagonal head self-drilling screws

6.3

Test results [kN] Sp-6.3-01

Sp-6.3-02

Sp-6.3-03

12.86

11.51

11.54

Average strength [kN]

Standard deviation

coefficient of variation

11.97

0.77

0.06

104

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107

Fig. 19. Frame-to-strap connection tests: a) CLE and CHE specimens; b) set-up and instrumentations.

Table 11 Results of frame-to-strap connection tests. Specimen type

Steel grade

Thickness [mm]

No. screws

Label

Ft [kN]

ke [kN/mm]

ke/na [kN/mm]

Failure modeb

CLE

S280GD

1.5

16

S280GD

2.0

24

85.50 85.34 85.12 85.32 0.19 0.00 171.35 169.73 169.94 170.34 0.88 0.01

48.01 43.89 40.21 44.04 3.90 0.09 63.08 73.12 66.39 67.53 5.12 0.08

3.00 2.74 2.51 2.75 0.24 0.09 2.63 3.05 2.77 2.81 0.21 0.08

NSF NSF NSF

CHE

CLE-01 CLE-02 CLE-03 AVc DEV.STd C.o.Ve: CHE-01 CHE-02 CHE-03 AVc DEV.STd C.o.Ve:

a b c d e

NSF NSF NSF

n: number of screws; NSF: net section failure of strap-bracing. AV: average of the experimental values. DEV.ST: standard deviation of the experimental values. C.o.V: coefficient of variation of the experimental values.

Table 10 provides the experimental results for a single tested screw resistant section and their average value. The experimental results are little scattered with a coefficient of variation b6%. 6.3. Tests on frame-to-strap connections The behaviour of CFS strap-braced stud walls is particularly affected by the design of frame-to-strap connections. For this reason, a series of

180 160

CLE

F [kN]

CHE

140 120 100 80 60 40 20 d [mm]

0 0

2

4

6

8

10

12

14

Fig. 20. Frame-to-strap connection tests - load vs. displacement curves.

tests on connection prototypes reproducing the joints between the gusset plate and strap-brace was carried out. In particular, two types of specimens were tested: CHE connections, adopted in the WHE walls, and CLE connections, adopted in the WLE walls (Fig. 19a). The connection specimen consisted of two steel plate connected to each other by means of 6.3 mm diameter hexagonal head self-drilling screws. CHE and CLE specimens were made with S280GD steel grade plates with thickness of 2.0 mm and 1.5 mm, respectively. The tests were performed by using a universal test machine and displacements at a rate of 0.05 mm/s. The data were recorded with a sampling frequency equal to 5 Hz. Two LVDTs, fixed by means of a steel support to the diagonal strap plate, measure the relative displacements between diagonal strap plate and gusset plates, through a contrast element fixed to gusset plate, whereas three strain-gauges placed at the centre and at the two strap ends recorded the strain of the diagonal strap (Fig. 19b). Table 11 lists the experimental values of the failure load (Ft), stiffness (ke) and stiffness for a single screw (ke/n, in which n is the screws number) together with the observed failure mode for each performed test. In addition, Table 11 provides the average experimental values, standard deviation and coefficient of variation (C.o.V.) for strength and stiffness. Fig. 20 shows the experimental response curves obtained for all tests, where d represents the displacement measured by the LVDT on connection side where the failure was observed. The coefficients of variation show that the experimental failure loads and stiffness were narrowly distributed, with C.o.V. always b 5% and 9%

L. Fiorino et al. / Journal of Constructional Steel Research 127 (2016) 92–107

CHE specimens

105

CLE specimens

Fig. 21. Frame-to-strap connection tests - failure mode (net section failure of strap-bracing).

respectively. For all tests the failure mechanism was the net section failure of the steel plate (Fig. 21). 6.4. Tests on hold-down devices The local response evaluation of the investigated X-braced CFS systems was completed with four tensile tests on hold-down devices. The two types of hold-down devices adopted for the two different configuration walls were tested: HD-H, adopted in the WHE wall, and HD-L, adopted in the WLE wall. The hold-down devices where made of a S235JR steel grade U-profile welded to a S275JR steel grade plate. In particular, the U-profile had a thickness of 8.0 mm for HD-H devices and 5.0 mm for HD-L devices; whereas the steel plate had a thickness of 20.0 mm for HD-H devices and 15.0 mm for HD-L devices. The hold down shear connections were made of four M18 or M14 bolts for HDH or HD-L devices, respectively, whereas the tension connection was made of one M24 or M20 bolts for HD-H or HD-L devices, respectively. All bolts were 8.8 bolt grade. Fig. 22 shows the drawing of the devices. The adopted test set-up is shown in Fig. 23. Single test was carried out on two devices connected back-to-back by to a 30.0 mm thick S355JR steel grade inner plate by means of hold-down shear connections. The inner plate was clamped to the upper wedge grip of the testing machine. Furthermore, each device was connected at the bottom

side to a stiffened T-shaped steel element by means of the hold-down tension connections. The T-shaped element was clamped to the bottom wedge grip of the testing machine. The tests were performed by subjecting the specimens to progressive displacements up to failure or maximum load capacity of the universal testing machine. The displacement-controlled test procedure involved displacements at a rate of 0.05 mm/s and data recorded with a sampling frequency of 5 Hz. Two LVDTs (L1 and L2) were used for measuring the relative displacements between the U-profile of the holddowns and T-shape element of the set-up. Table 12 shows a summary of experimental results, whereas Fig. 24 shows the experimental force-displacement curves. The limit resistance (Flim) of the device was defined as the force corresponding to a limit displacement (dlim), set equal to 4.7 mm, according to the AISI S913 [29]. In addition the stiffness (ke,d(lim)) was defined as the secant stiffness at a limit displacement (dlim). Tests of HD-H specimens were carried out up to reach the maximum load capacity of the universal testing machine (500kN). These specimens showed only a strong deformation of the base plate for high values of the applied load, without any rupture (Fig. 25a). However, the limit resistance (Flim) was reached. HD-L specimens showed two different collapse mechanisms. In particular, during the first test (HD-L01) a tensile failure of the CFS U-profile was observed (Fig. 25b),

Fig. 22. Hold-down specimens.

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Fig. 23. Hold-down tests - set-up.

whereas the second test (HD-L-02) was characterized by the bending failure of the base plate (Fig.25c).

7. Conclusions and further developments The current paper presents and discusses an experimental investigation carried out for characterizing the seismic behaviour of CFS strapbraced stud walls. The experimental campaign is principally focused to study the global response of CFS strap-braced walls designed according to an elastic approach. Moreover, the local behaviour is investigated by tests on material and main connection systems (material, screws, frame-to-strap connections and hold-down devices). A good correspondence between the experimental and theoretical expected results is highlighted, particularly for the wall strength. The experimental study also shows that the wall corners should be carefully designed, since their behaviour could significantly affect the overall wall response. The tested walls exhibited the typical response of structural elements designed without capacity design criteria. Indeed, their inelastic behaviour can be affected to the effect of non-ductile phenomena, as

the gusset-to-track connection failure and combined bending and compression axial load failure of the chord studs. In addition, for the investigated cases were not observed appreciable degradation effects of the cyclic action for the walls tested with the adopted loading protocol. As further developments, these experimental results could be considered as a reference for numerical studies aimed to define seismic design criteria for the investigated systems. In addition, a shake table test on a reduced-scale three-storeys two-bays 3D structure, representative of study case illustrated in Section 2, will be carried out for obtaining a complete overview of the seismic performance of the investigated structural systems.

Acknowledgments The authors acknowledge the Lamieredil S.p.A. POR CALABRIA FERS 2007-2013 (CC N° 2007 IT 161 PO 008) for the financial support to the research activity.

300 Table 12 Results of hold-down tests. Label

Flim [kN]

ke,d(lim) [kN/mm]

Failure modea

HD-L

HD-L-01 HD-L-02 AVb DEV.STc C.o.V.d HD-H-01 HD-H-02 AVb DEV.STc C.o.V.d

170.4 110.8 140.6 42.2 0.30 226.8 254.7 240.7 19.7 0.08

36.3 23.6 29.9 9.0 0.30 48.3 54.2 51.3 4.2 0.08

T B

HD-H

a b c d

ke,d(lim)

250

Specimen type

HD-H-01 HD-H-02 HD-L-01 HD-L-02

F [kN]

200 150

– –

T: tension failure of the U-shaped element; B: bending failure of the base plate. AV: average of the experimental values. DEV.ST: standard deviation of the experimental values. C.o.V: coefficient of variation of the experimental values.

100 50 dlim=4.7mm

0 0

5

10

d [mm] 15

20

25

Fig. 24. Hold-down tests - load vs. displacement curves.

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Fig. 25. Hold-down tests - failure modes: a) strong deformation of the base plate; b) tensile failure of the U-profile; c) bending failure of the base plate.

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