cyclic behaviour of a full scale rc structural wall

Engineering Structures 25 (2003) 835–845www.elsevier.com/locate/engstruct

Cyclic behaviour of a full scale RC structural wall

P. Riva∗, A. Meda, E. GiurianiCivil Engineering Department, University of Brescia, Via Branze, 38, I-25123 Brescia, Italy

Received 8 August 2002; received in revised form 7 January 2003; accepted 7 January 2003

Abstract

The results of an experimental test on a full scale RC structural wall subjected to cyclic loading are herein presented. The testedspecimen is representative of a wall in a four storey building with one underground floor, designed for moderate seismic actions(PGA=0.20 g) adopting the European Seismic Code (Eurocode 8, EC8). The experimental specimen is 15.5 m long and has atransverse section of 2800×300 mm. The boundary conditions consist of simple supports at the foundation and ground floor levels.The wall behaviour has been studied both under service conditions, up to yielding, defined as the point to which first yield of theoutermost rebars corresponds, and ultimate conditions, up to collapse. 2003 Elsevier Science Ltd. All rights reserved.

Keywords: RC structural wall; Seismic design; Cyclic loading; Full scale test

1. Introduction

In seismic zones, building resistance to earthquakes isoften ensured by adopting structural systems where seis-mic actions are assigned to structural walls, designed forhorizontal forces and gravity loads, while columns andbeams are designed only for gravity loads [1,2]. Thesesystems, being stiffer than earthquake resisting frames,allow a better displacement control, limiting damage ininternal partition walls and non structural elements. Onthe contrary, frame structures generally exhibit greaterductility, at the expense of large displacements and inter-action problems between structural and non-structuralelements.

Extensive experimental results concerning the behav-iour of walls of different slenderness ratio subjected tovarious loading conditions are available in the literature(e.g. [3–6]). These tests are generally limited to smallscale specimens, typically from 1:2 to 1:3 scale, whileexperimental evidence on the behaviour of full-scalewalls is presently scarce. The results have shown thatthe inelastic response of slender walls, characterized byheight-over-width ratios larger or equal to 2, is con-

∗ Corresponding author. Tel.:+39-030-3715502; fax:+39-030-3715503.

E-mail address: [email protected] (P. Riva).

0141-0296/03/$ - see front matter 2003 Elsevier Science Ltd. All rights reserved.doi:10.1016/S0141-0296(03)00020-8

trolled by flexural deformations in a plastic hinge at thebase of the wall.

To achieve adequate ductility, an essential role isplayed by confining steel, placed at the edge of the sec-tion in order to control concrete crushing and longitudi-nal reinforcement buckling. Shear strength is providedby distributed vertical and horizontal reinforcement onboth wall faces. Inclined reinforcement is sometimesneeded for protection against sliding shear.

The scope of the present research is to partially fill inthe gap concerning full scale tests on slender shear walls,by analysing the behaviour of a full size structural wallunder cyclic loads, with particular attention to its duc-tility, dissipated energy, damage progression, andresisting mechanisms. The tested specimen is representa-tive of a shear wall of a four storey building, with oneunderground storey and a box foundation system (Fig.1). In box foundation systems (Fig. 2), horizontal forcesare resisted by the diaphragm actions of ground andbasement level slabs, leading to a considerable reductionin the wall foundation dimensions [7]. In this case, thecritical section of the wall is located at the ground level,where bending actions are predominant, whereas theunderground part of the wall exhibits a typical shearpanel behaviour. Due to expected high shear forces inthe underground part of the wall, its thickness isincreased.

The building, hence the experimental wall, was

836 P. Riva et al. / Engineering Structures 25 (2003) 835–845

Nomenclature

Agt steel elongation at 0.99 Rm after the peak, measured over five bar diameter;F net applied force at each jack, equal to the difference between the total jack force and the force

necessary to obtain a bending moment equal to 0 at the critical section;PGA peak ground accelerationq structural coefficient according to EC 8 (=3 for class M structural walls);Rcm mean cube concrete strength;Re reinforcing steel yield strength;Rm reinforcing steel ultimate strength;M bending moment at the critical section;V shear force at the critical section;d displacement at the wall end;gRd overstrength coefficient according to EC8 (=1.15 for class M structural walls);gc, gs partial safety factors for concrete and steel, respectively;

Subscripts

n nominal ultimate values;Sd acting design values;Rd resisting design values;u ultimate values at collapse;y experimental yield values;yt values corresponding to the theoretical first yield of the critical section.

designed according to Eurocode 8 (EC8) [8–10],assuming Medium ductility class (structural coefficientq=3), a peak ground acceleration PGA=0.20 g, typicalfor medium seismicity zones, and a Soil type B(equivalent to Soil type C in the latest EC8 version [13]).

2. Structural wall description and test set-up

The structural wall was designed according to EC8[8–10], assuming Medium ductility class (structuralcoefficient q=3) and a peak ground accelerationPGA=0.20 g, typical for medium seismicity zones. Ahigher ground acceleration could not be adopted, due tolimitations in the testing loading frame available. Veri-fication of sectional strength was carried out accordingto Eurocode 8 and Eurocode 2 [8–11].

Fig. 3 illustrates the wall dimensions and steelreinforcement detailing. The wall dimensions were: sec-tion 2800×300 mm outside of the supports,2800×400 mm between the supports, length outside thesupports 12.5 m, and 16 m total length. At the groundand basement levels, two ribs were inserted to simulatethe floor diaphragms.

As prescribed by EC8, the main flexural reinforcementwas concentrated in two chords at the edges of the wall,where the reinforcement was heavily confined. Design

shear force VSd at the critical section was determinedbased on the overstrength prescribed by EC 8 [10] as:

VSd � e·V’Sd � q��gRd

q·MRd

MSd�2

� 0.1�Se(TC)Se(T1)

�2

·V’Sd

�1.50·V’Sd,

where: q=3 is the structural coefficient, gRd=1.15 is theoverstrength factor, MRd and MSd are the resisting anddesign bending moment, respectively, Se(TC)/Se(T1) isthe ratio between the elastic response spectrum ordinatesat the end of the constant acceleration branch and at thefundamental period, respectively, while V’Sd is the shearforce derived from the analysis.

Shear reinforcement was limited to vertical and hori-zontal bars. No inclined reinforcement was inserted, asthe theoretical strength evaluated according to the codewas in excess of the design shear force. The governingshear resisting mechanism was found to be sliding shearat the critical section, expressed as:

VRd,s � Vdd � Vid � Vfd,

where Vdd is due to dowel action of the web reinforce-ment across the critical section, Vid is due to inclinedreinforcement (=0 in the experimental specimen), Vfd isdue to friction effects in the compression chord, whichis the predominant resisting term.

837P. Riva et al. / Engineering Structures 25 (2003) 835–845

Fig. 1. Four-storey building adopted for the structural wall design.

Fig. 2. Foundation box system for the wall.

Table 1 shows the design bending moment and shearstrengths of the critical section computed considering thematerial safety factors (gc=1.5, gs=1.15) and the designstrength of concrete (C30/37) and steel (B500B) (MRd,VRd), and the nominal strength values determined bymeans of experimental tests on the materials (Mn, Vn).The experimental bending moment and shear force atstructural yield (My, Vy) and at collapse (Mu, Vu) are alsoreported in the same table. It is observed that due to thetest set-up, no axial force is present in the wall. As aconsequence, the yield and ultimate moment as well asthe shear strength of the specimen are smaller than thosein a real structural wall.

The average material characteristics as determined onconcrete cube specimens (150×150×150 mm) and steelbars were:

� Concrete average cube strength: Rcm=40.7 MPa;� Reinforcing steel yield strength: Re=560 MPa;� Reinforcing steel ultimate strength: Rm=640 MPa;� Reinforcing steel elongation at 0.99 Rm after the

peak: Agt=8.4%.

Being the reaction structure available, a prestressed con-crete underground caisson which allows testing struc-tures of span up to 40 m, the wall had to be placed hori-zontally, keeping the axis of maximum inertia vertical,as shown in Fig. 4.

The wall was placed on two RC supports, (a) inFig. 4, aligned with the ribs simulating the ground andbasement floor diaphragms, (b) in Fig. 4, and fixed tothe caisson by adopting post tensioned 0.6” strands andhigh strength �32 bars, (c) in Fig. 4. Strands and barspost tension was such that no decompression of the sup-port would occur during testing.

Two steel frames, (d) in Fig. 4, were placed near theloading positions in order to avoid lateral instability. Thesafety of the system was improved by inserting a sup-plementary frame between the two supports, (e) in Fig.4, which would intervene whenever a lack in post-ten-sion would induce a support decompression.

The loads were applied at two points by means ofhydraulic jacks. The position of the jacks was definedto obtain the same bending moment and shear forcearound the critical section as the one resulting from theanalysis of the four storey building (Fig. 5). Moreover,the load position was such that the same force could beapplied, greatly simplifying load control. In order toapply cyclic reverse loads, four jacks were adopted, twoacting upward, (a) in Fig. 6, and two downward, (b) inFig. 6. The jacks acting upward were placed betweenthe wall and the loading bench, while those acting down-ward were placed in two windows opened in the walland connected to the caisson with two high strength �32bars. The position of the opening was such that the jack


Fig. 3. Wall dimension and steel reinforcement.

Table 1Theoretical and experimental bending moment and shear strengthvalues: design (MRd, VRd,s); nominal (Mn, Vn); structural yield (My, Vy);collapse (Mu, Vu)

MRd=4015 kN m VRd,s=715 kNMn=5910 kN m Vn=1165 kNMy=5300 kN m Vy=615 kNMu=6200 kN m Vu=720 kN

Fig. 4. Test set-up: wall supports (a); ribs simulating the ground and basement floor diaphragms (b); post tensioned strands and bars (c); steelframe to avoid lateral instability (d); additional frames for improving safety in the test set-up (e).

would act along the wall neutral axis, thus limiting theirhorizontal displacement.

The applied load was measured by means of a fullbridge resistive pressure transducer placed on the pumpmanifold. The displacements were measured using 17potentiometric transducers as shown in Fig. 6: two wiretransducers (16, 17) measured the vertical displacementof the wall; 11 linear transducers (1–8, 14, 15) measuredthe displacements in the upper and lower chords closeto the critical section; two linear transducers (11, 12)


Fig. 5. Action diagram: comparison between true scheme and the two jacks scheme.

Fig. 6. Jack positions (upward jacks (a) and downward jacks (b)),and measurement devices.

were used for monitoring the deformation of the panelbetween the supports; two linear transducers (9, 10)measured the displacement between the wall and thecaisson at the supports in order to monitor any potentialsupport decompression. All the signals were conditionedby adopting a data acquisition system (Mod. UPM 100by HBM) and recorded on a PC.

At first, the test was performed by applying cyclicloads with increasing amplitude (Fig. 7), until the theor-etical yield load Fyt (i.e. theoretical load for which thefirst yield (Re=560 MPa) of the outermost rebar occurs

Fig. 7. Loading history up to the theoretical yield load: M/Myt versuscycle number.

at the critical section), corresponding to an experimentalmaximum displacement at the wall end (dyt) equal to90 mm, was reached. After three cycles with maximumdisplacement equal to dyt, the load was further increasedin both directions until the displacements correspondingto the structural yield in both directions were detected(+dy�120 mm and �dy��120 mm). These displace-ments were determined by intersecting two lines tangentto the load–displacement experimental curve in the IIstage (after cracking) and in the III stage (after yielding).The following loading history was defined by imposingcycles of increasing displacement amplitude until failurewas reached (Fig. 8). The collapse occurred during thesecond cycle at a displacement equal to three times thestructural yield displacement, also equal to four times thetheoretical first yield displacement (3dy�4dyt�360 mm).


-4

-3

-2

-1

0

1

2

3

4

24 25 26 27 28 29 30 31 32

cycles

δ/δ

y

Fig. 8. Loading history after the structural yield load: d/dy versuscycle number.

3. Experimental results

Fig. 9 shows the load versus end displacement curves(F–d) for the whole loading history, where F is the nettransverse load, equal to the difference between the totaljack load and the jack load necessary to equilibrate thewall weight, i.e. the load which annuls the bendingmoment in the critical section. The results are discussedin the following, by analysing separately the load cyclesbefore yielding, typical of service conditions, and afteryielding, typical of design seismic loads. The results ofthe experimental test not reported and discussed hereinmay be found in [12].

3.1. Behaviour under service loads

The behaviour in service conditions was analysed byapplying load cycles lower or equal to the first theoreti-cal yield load, defined as the load for which the stressin the external reinforcement at the critical section isequal to Re=560 MPa. In the present case, the theoreticalyield bending moment at the critical section and applied

Fig. 9. Force F versus end displacement d curve for all the cycles.

Fig. 10. Force F versus end displacement d for the cycles up to dyt.

load are equal to Myt=4851 kNm and Fyt=268 kN,respectively, while the corresponding experimental enddisplacement is equal to dyt�90 mm.

Fig. 10 shows the F–d curve for the cycles up to dyt.From this figure, the following observations are made:

� due to cracking development, the wall stiffnessdecreases for increasing displacement amplitudes;

� cycles with constant amplitude show an almost con-stant dissipated energy, with reduced accumulateddamage after each cycle (Table 2);

� the wall cyclic behaviour is nearly linear elastic, as

Table 2Dissipated energy for cycles up to dyt

Cycle d+–d� F+–F� Dissipatedenergy percycle

mm kN J

6 26.0 203 5307 26.1 203 5658 26.5 203 557

9 26.4 191 34110 25.2 190 33011 28.2 190 434

12 30.0 208 506

13 61.6 298 1 433

14 71.9 331 1 66715 77.3 345 2 23616 77.7 343 2 06417 78.5 339 1 878

18 100.3 371 2 20319 100.4 369 1 998

20 179.3 536 9 83921 198.5 545 9 74122 199.7 561 7 572


confirmed by comparing the experimental responsewith the theoretical elastic behaviour of a crackedconcrete wall (heavy line in Fig. 10).

The main cracks in the zone close to the critical sec-tion (Fig. 11) have limited inclination, proving that thebehaviour is governed by bending. The crack distance isclose to the stirrup spacing in the external chords, wherelongitudinal reinforcement is concentrated, while in themiddle part of the critical section the cracks tend tomerge in a lower number of cracks, characterised bylarger opening and greater inclination, thus showing theinfluence of shear stress. In any case, crack opening islimited (�0.8 mm for the cycle at dyt) and an overallgood service behaviour of the wall is observed.

As for the zone between the supports, a diffused crackpattern was detected, with a crack inclination of about45°, typical of panels loaded by pure shear, and areduced crack opening (lower than 0.1 mm). It isremarked once more that in order to improve the shearstrength in a zone where shear is particularly severe,both the wall thickness (400 mm) and the stirrups(�12@200 mm) were increased in this area.

3.2. Behaviour under ultimate loads

The experimental behaviour following the theoreticalyield is presented in Fig. 12 (load versus end displace-ment F–d) and in Fig. 13 (critical section bendingmoment versus end displacement M–d). These figuresclearly show the structural yield point, corresponding toa wall end displacement approximately equal to dy=120 mm, with bending moment My and shear force Vy

given in Table 1. Furthermore, it is possible to observethat due to the different position of the jacks in the uplift(Fig. 6(a)) and lowering phases (Fig. 6(b)), the load dis-placement response (F–d) is non-symmetrical, while themoment-displacement one (M–d) is symmetrical.

The experimental curves show a sudden decrease

Fig. 11. Crack pattern at the theoretical yielding (principal cracks aremarked).

Fig. 12. Force F versus end displacement d for the cycles after dy.

Fig. 13. Bending moment at the critical section M versus end dis-placement d for the cycles up to 3 dy.

upon load reversal, due to the sudden pressure releasein the single effect jacks when the retaining valve isopened.

The test results allow the following observations:

� Cyclic response is stable up to failure and no relevantpinching is observed in the cycles;

� The dissipated energy for each cycle, reported inTable 3 and Fig. 14, increases almost linearly with

Table 3Dissipated energy for cycles beyond dy

Cycle d+–d� F+–F� Dissipatedenergy percycle

mm kN J

23 242.1 575 16 19924 244.4 594 17 75225 243.2 553 17 148

26 342.9 735 43 852

27 480.3 747 125 381

28 600.4 764 209 745

29 724.2 746 303 738


Fig. 14. Dissipated energy per cycle.

the displacement amplitude, while it remains almostconstant for repeated cycles at constant amplitude(cycles 23, 24, 25), thus showing that damage pro-gresses in a stable manner.

Fig. 15 illustrates the average strain in the chords dur-ing the cycles. This figure shows that bending reinforce-ment progressively yields with increasing cycle ampli-tude. At collapse, all of the reinforcement within theinstrumented zone (up to 1600 mm from the basesection) may be considered as yielded.

Fig. 16 shows the development of the crack pattern inthe critical section for increasing cycle amplitude. Image16(a), concerning the last cycle at yield (cycle 25), illus-trates that the main cracks are evident but their openingis still small. More evident are the cracks in the cycleat 2 dy (Fig. 16(b) and (c), cycle 27). In detail, it ispossible to observe a vertical crack close to the ‘groundfloor diaphragm’ (section close to the support) and a ser-ies of wide opened and curved cracks, with the largeststarting approximately 400 mm from the support andreaching the diaphragm section at middle wall depth. At2 dy, a distributed crack pattern with limited opening isobserved in the chords. The cracks tend to merge in themiddle part of the wall in the aforementioned curvedcracks. Starting from the cycle at 2.5 dy (Fig. 16(d) and(e), cycle 28) also the cracks in the chords exhibit awider opening, while the main curved crack, whicheventually will lead to the wall failure, shows a consider-ably increased opening.

In the cycles at 3 dy (Fig. 16(f) and (g), cycles 29 and30) the damage is relevant, especially in the zonebetween the chords. By comparing the crack pattern cor-responding to the uplift (Fig. 16(f)) and lowering(Fig. 16(g)) phases, it is possible to note that in the lattercase the dead load, increasing the bending momentgradient, induces a higher crack localisation with widercrack opening.

Fig. 16(h) shows the wall after the uplift phase of thesecond cycle at 3 dy, just before failure, which occurred

in the following unloading phase. The main crack exhib-its a maximum opening of approximately 50 mmtowards the wall mid-depth and 10 mm at the chords.In the chords, concrete spalling, due both to compressionforces and rebar bending, is observed. As shown inFig. 16(i), the large crack opening in the middle partof the wall leads to a marked strain localisation in thelongitudinal shear reinforcement, resulting in its tensilefailure with necking. Concerning shear resistance, thewide crack opening observed is not compatible with anyaggregate interlock effect. Furthermore, the subsequentfailure of the longitudinal panel reinforcement led to aconsiderable reduction in the shear strength for dowelaction. As a consequence a shear failure occurred duringthe unloading phase, when the beneficial effect of com-pression in the upper chord, which enables the shearstrength contribution due to friction, ceased to exist(Fig. 16(l)). Fig. 16(m) shows that the main rebars arebent due to dowel action and consequent vertical walldisplacement.

It is important to observe that with the wall beingplaced horizontally on the loading bench, the beneficialeffect of the axial force due to gravity loads is notpresent in the tested wall, thus leading to an anticipatedshear collapse.

Finally, little damage is present in the zone betweenthe two supports (Fig. 16(n)), where the observed crackpattern confirms a shear panel behaviour.

The wall ductility anyway ensures adequate behaviourunder the design earthquake actions. In fact, themaximum obtained ductility is equal to 4dyt, larger thanthe assumed design ductility, approximately equal toqdyt=3dyt. Furthermore, it is observed that a very largeultimate displacement was obtained, equal to 360 mmor l/35, l=12.50 m being the wall height.

Nonetheless, it is important to note that collapse didnot occur in bending, but was due to shear as a conse-quence of a lack of longitudinal reinforcement in thewall web (between the chords). In fact, the amount ofweb reinforcement provided in accordance to EC8 [10]was not enough to limit the observed crack opening, andfriction contribution to shear strength resulted in beingmuch smaller than expected. Accordingly, a higheramount of reinforcement should be adopted in the webto ensure a greater contribution of dowel action and ofinclined bars, whenever present. This could be obtainedby either reducing the friction contribution, or increasingthe overstrength factor when determining the designshear force.

3.3. Conclusions

This paper presents the results of an experimental testconducted on a full scale RC structural wall subjectedto cyclic reverse loading with amplitude increasing up tocollapse. The results allow the following considerations:


Fig. 15. Average strain in the chords along the critical zone.

� service behaviour was found to be mostly linear elas-tic, with reduced damage and small dissipated energy;

� the behaviour after yielding showed a progressivedamage of the wall with increasing imposed displace-ment cycle amplitude. The damage is mainly localisedat the critical section, where cracking progressivelydeveloped into large, wide open cracks, and concretecrushing was observed;

� no significant strength and/or stiffness degradationwas observed during cycles following yield;

� the collapse mechanism was governed by shear, withformation of a single large crack near the base sec-tion, almost parallel to the ground floor diaphragm,leading to the tensile (necking) failure of the longi-tudinal wall reinforcement;

� a considerable ultimate displacement was obtained,

equal to 360 mm or l/35, l=12.50 m being the wallheight;

� regardless of the early shear failure, the ductility coef-ficient with respect to the theoretical first yield,defined as the instant when the outermost rebars reachthe yielding stress, was equal to du/dyt=4, while theductility coefficient with respect to the observed struc-tural yield was equal to du/dy=3;

� the stability of the wall response up to collapse leadsto the conclusion that a considerable ductility marginwas still available with respect to bending failure;

� the shear reinforcement provided at the critical sectionproved to be insufficient to avoid an early shear fail-ure;

� the latest version of EC8 [13] prescribes a minimumoverstrength for shear design equal to 1.50 for duc-


Fig. 16. Cracking development around critical section from dy to collapse.

tility class M walls, while for class H structures thesame overstrength factor adopted by the 1994 EC 8version [10] is given. Concerning sliding shear fail-ure, the same formulation adopted in [10] is proposed,with a slight increase of the friction contribution fornormal strength concrete. Based on the results hereinpresented, the adopted approach might lead to non-conservative design for sliding shear effects, as shearfriction resistance might be overestimated;

� in order to prevent sliding shear failure, particularlywhen axial force in the wall is negligible, the webreinforcement amount should be increased. This couldbe obtained by either reducing the friction contri-

bution, or increasing the overstrength factor whendetermining the design shear force.

Acknowledgements

The research was co-financed by MURST (ItalianMinistry of University, Scientific and TechnologicalResearch) within the program COFIN-99 “Safety of RCstructures under seismic actions with reference to designcriteria of ultimate strength and damage limitation givenby EC8” . The contribution of UNIECO s.c.r.l., Cal-cestruzzi s.p.a., Ferriera Valsabbia s.p.a., Italcables s.p.a.


towards the construction of the experimental specimenis kindly acknowledged. Finally, the authors are gratefulto Mr GianPaolo Beccari, Mr Giovanni Grazioli, MrMaurizio Lancini, Mr Denny Rivetti, Mr Marco Sand-rini, Mr Gabriele Tosi, who carried out both design andexperimental test of the wall under their respective Mas-ter Thesis projects.

References

[1] Paulay T, Priestley MJN. Seismic design of reinforced concreteand masonry buildings. New York: J.Wiley & Sons, 1992.

[2] Paulay T. Earthquake-resisting shearwalls—New Zealand designtrends. ACI Journal 1980;77(3):144–52.

[3] Bertero VV, Popov EP, Wang TY, Vallenas J. Seismic designimplications of hysteretic behavior of RC structural walls. In:Proceedings of the 6th World Conference on Earthquake Engin-eering, New Delhi, vol.5. 1977. p. 159–65.

[4] Paulay T, Priestley MJN. Stability of ductile structural walls. ACIStructural Journal 1993;77(4):385–92.

[5] Pilakoutas K, Elnashai AS. Cyclic behaviour of RC cantilever

walls, Part I: Experimental results. ACI Structural Journal1995;92(3):271–81.

[6] Tasnimi AA. Strength and deformation of mid-rise shear wallsunder load reversal. Engineering Structures 2000;22:311–22.

[7] Giuriani E, Gubana A, Chiari G, Contessi R. Box foundation forstructural walls under seismic actions (in Italian). TechnicalReport, University of Brescia, Civil Engineering department,1996.

[8] Eurocode 8, “Design provisions for earthquake resistance ofstructures—Part 1–1: General rules, seismic actions and generalrequirements for structures” , ENV 1998-1-1, 1994.

[9] Eurocode 8, “Design provisions for earthquake resistance ofstructures—Part 1–2: General rules, general rules for buildings” ,ENV 1998-1-2, 1994.

[10] Eurocode 8, “Design provisions for earthquake resistance ofstructures—Part 1–3: General rules, specific rules for variousmaterials and elements” , ENV 1998-1-3, 1995.

[11] EUROCODE 2, “Design of concrete structures—Part 1–1: Gen-eral rules and rules for buildings” , ENV 1992-1-1, 1991.

[12] Riva P, Meda A, Giuriani E, et al. Cyclic behaviour of a fullscale RC structural wall (in Italian). Technical Report N.18/2002,Dip. di Ingegneria Civile, Universita di Brescia, November 2002.

[13] EUROCODE 8: “Design of structures for earthquake resistance—Part 1: General rules, seismic actions and rules for buildings”PrEN 1998-1-1, DRAFT No. 5, Revised Final Project Team Draft(preStage 49), Doc CEN/TC250/SC8/N317, May 2002.

Documents

cyclic behaviour of a full scale rc structural wall