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NEW GENERATION OF EARTHQUAKE RESISTING SYSTEMS

Jos I. RestrepoAssociate Professor

University of California, San Diego

UNITED STATES

Keywords: energy dissipation devices, precast concrete, prestressed concrete, self-centering sys-tems

1 INTRODUCTION

The impact and cost of the consequences of damage caused by earthquakes worldwide duringthe past two decades have raised serious questions of whether current building seismic design phi-losophies are satisfying the needs of modern society. Most seismic design standards are based on alife-prevention approach where building structural and non-structural damage is accepted providingthat collapse is avoided. No other economic parameters, such as the cost of damage to equipmentand stored goods and the cost associated with loss of operation following a moderate/strong earth-quake, are currently accounted for in the design process. The rapid advance in technology has meantthat often, in many countries, the cost of equipment, of stock and the loss of business operation in theaftermath of a moderately or strong earthquake are higher than that of the building itself. As a result,and in spite of efforts aimed at improving seismic design methodologies and construction practices,economic losses due to earthquakes have increased rather than decreased in the past decades.

The framework for a performance-based design was first proposed in an attempt to encompass,directly or indirectly, all parameters within a set of objectives aimed at ensuring predictable behavior ofthe entire building [1] and to, ultimately, mitigate economic losses. Within the framework of perform-ance-based seismic design, significant effort has been made in recent years at developing new meth-odologies [2, 3] and new structural systems in a way that the design objectives could be achieved.

This paper describes a number of emerging structural wall systems aimed at minimizing seismi-cally induced structural damage. The design of the system within a displacement-based seismic de-

sign procedure is also discussed herein. Finally, the paper describes the results of experimental workperformed on these systems.

2 NEED FOR DEVELOPMENT OF NEW STRUCTURAL SYSTEMS

The cost associated with the loss of business operation, damage to equipment and structuraldamage following a moderately strong earthquake can be significant to modern society, particularly inthose centers of advanced technology. Such cost is often comparable, if not greater, to the cost of thebuilding itself. With current design approaches, most building structures are designed to respond be-yond the elastic limit and eventually to develop a mechanism involving a number of flexural plastichinges. For this reason plastic hinges are especially detailed for ductility and for energy dissipation.

Whilst the concept of inelastic response is very appealing, it has two main drawbacks: First, re-

gions in the principal lateral force resisting system will be sacrificed in moderately strong earthquakesand in need of repair, or damaged beyond repair in strong earthquakes. Second, current design ap-proaches are based on the premise that large energy dissipation capacity is necessary to mitigate theeffects induced by earthquakes. This premise has very often led to the notion that a good structuralsystem should be characterised by "fat" hysteresis loops. As a very large fraction of the input energyis expected to be dissipated by hysteresis due to flexural yielding in the plastic hinges, significant re-sidual displacements could be expected in a building after an earthquake. In a parametric analysis onsingle degree of freedom oscillators, Christopoulos et al. [5] noted that oscillators characterised by aflag-shaped hysteretic response, which have significantly less energy dissipated per cycle than elasto-plastic oscillators, were subjected in general to lower ductility demands than elasto-plastic oscillatorswith the same level of strength. An added benefit of a flag-shaped hysteretic response was the lack ofresidual displacements.The use of some widely accepted structural systems can become questionable in light of post-

earthquake occupancy and reparability. For example, it has long been recognized that, in momentresisting frames of high-rise buildings, the best mechanism of inelastic deformation is the beam

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Seismic design of concrete structuresSession 6


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sidesway mechanism, see Fig. 1. This mecha-nism distributes the plastic rotations throughmost beams in the frame, precluding the concen-tration of rotation in any particular location.Nevertheless, the distribution of plasticity in somany locations in the structure can have a

severe impact on the cost of repair, particularly ifa prefabricated faade or ceiling hide suchregions. This potential problem is more relevantto current design practice as little emphasis isgiven to post-earthquake reparability issues, likeproviding suitable access for grouting residualcracks in cladded perimeter frames. Suchproblems were observed in Kobe following the1995 Hyogo-ken Nanbu earthquake [4].

Another important issue in seismic designnowadays deals with current societal perform-ance expectations. While the principle of mitigat-

ing loss of life in a strong earthquake still prevails,society expects buildings to survive a moderatelystrong earthquake with no disturbance to busi-ness operation. This implies that repairs of in-cipient structural damage may not be economical,and may no longer be tolerated in small andmoderately strong events. Alternatives such asseismic isolation provide some of the desirableresponse characteristics, but the cost confinesuch alternatives to only landmarks of significantimportance and to bridge structures.

The main objective of this paper is to de-scribe some structural systems that possess self-

centering characteristics and are economicallyviable alternatives to current lateral force resist-ing systems. Particular emphasis is given to theresponse characteristics of cantilever walls thatare explicitly designed to self-center after an ex-cursion beyond the limit of proportionality. Thissystem encompasses a range of walls that arejointed at their base and are properly detailed torock about their foundation. These walls can beprestressed with unbonded tendons and can in-corporate devices designed especially for energydissipation. Jointed walls are grouped into rock-

ing or hybrid types depending on the lateralforce-lateral displacement response characteris-tics, see Fig. 2. Rocking walls show a non-linearelastic response, whereas hybrid walls show aflag-shaped response.

3 PRACTICAL DEVELOPMENT

In 1993, Priestley and Tao [6] proposed theuse of moment resisting frame systemsprestressed with partially unbonded tendons asthe primary lateral force resisting systems in

seismically prone areas. This proposal was sup-ported by a series of non-linear dynamic time-history analyses that showed the viability of such

Fig. 1 Typical beam sidesway mechanismin a ductile frame

Fig. 2Classification of jointed walls

Proceedings of the 1st fib Congress

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tailing to prevent the concrete from crushing prematurely.In the initial stages of lateral loading there is no contribution to the lateral displacement from de-

formations in the joint at the base of the wall, the wall panel deforming primarily inflexure. At point a in Fig. 5, the extreme fibre of the joint located at the base of the wall is decom-pressed. Further loading results in a gradual opening of the joint, and as a result, in a gradual reduc-tion in stiffness, see the dotted line ab in Fig. 5 When the gap in the joint is sufficiently wide, the en-

ergy dissipators attain the elastic limit (see point b in Fig. 5). Following this point the stiffness of thewall is further reduced. The stiffness of the system at loading beyond point b is related to the strain-hardening stiffness in the energy dissipators plus the stiffness of the tendons. At point cof Fig. 5, theenergy dissipators reach the plateau associated with the behaviour of steel beyond one-half the ulti-mate tensile strain. Hence, the remaining stiffness left in the system is solely due to the presence ofthe tendons that remain elastic. At point dof Fig. 5, the energy dissipators reach the limiting strainand are expected to fracture at any displacement beyond this point. The quadri-linear representationgiven by the lines oa, ab, bc, and cd is able to provide a good representation of the monotonic re-sponse of hybrid walls.

In a hybrid cantilever wall of rectangular cross section subjected to a single lateral force, the lat-eral force, Hr, causing decompression at the extreme fibre at one of the lower corners of the wall canbe determined from elastic theory. In this case the presence of reinforcing steel in the wall and of the

energy dissipators is ignored for practical purposes. Thus:

( )eff

wr

h

lNPH

6+=

(1)

where P and N are the prestressing force after losses and gravity load actions at the base of the wallpanel, respectively, and heffand l ware the distance from the base to the lateral force and the length ofthe wall, respectively.

The lateral displacement, r, at the point of application of the lateral force H r, is obtained by com-bining the flexural and shear deformations occurring within the wall panel. Using elastic theory, as-suming the wall panel is uncracked and that the shear modulus of concrete G c= 0.4E c, the followingexpression is obtained for r,

a

b

cdHu

Hp

Hy

Hr

r y p u

1

1

11

1

1

Kw

Kr

Kp

K + Ked p

K + Ked p

Kp

Lateral displacement

Partially-Prestresse

Prest ress ed-only

Ro cking-only

Lateralload

Fig. 5 Monotonic lateral force lateral displacement response of jointed systems [14]


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r

gc

eff

eff

wr H

IE

h

h

l

34

31

32

+= (2)

where Ec is the concrete Youngs modulus and Igis the second moment of area of the gross section.

When the wall is loaded towards point b, see Fig. 5, the neutral axis depth migrates towards theextreme fibre in compression. The resulting compressive force at the base of the wall approaches theextreme compressive fibre. The tangential stiffness Ked, Ked and K pof the lines depicted in Fig. 5 canbe determined using the mechanical model shown in Fig. 6. At ultimate load, the resulting force will bevery close to the extreme fibre in compression. If the position of the resulting force is assumed to re-main constant when loading through points bto d, the wall can be represented by the superposition ofa prestressed-only and a energy dissipation-only mechanisms, see Fig. 5. If gravity loading is smalland is ignored for practical purposes, Kr = 0 in Fig. 5. For a wall prestressed with tendons actingthrough the walls centerline and incorporating axial energy dissipation devices acting also at this loca-tion the stiffness Ked, Ked and K p are:

2

41

=

eff

w

ed

sdsed

hal

LAEK

(3)

2

'

4

1'

=

eff

w

ed

sds

edh

al

L

AEK

(4)

and

2

4

1

=

eff

w

sp

spsp

ph

al

L

AEK

(5)

where a/2 is the position of the resultant compressive force at the base of the wall, measured from theextreme compressive fibre; Es is the elastic modulus of steel in the axial energy dissipation devices;Es is the strain hardening modulus of steel in the idealised tri-linear stress-strain relationship shown inFig. 7; Asdis the area of a single energy dissipator; E spis the elastic modulus of the tendons; A spis thetotal area of the tendons and Lsdand L edare the tendons unbonded length and the yielding length ofthe axial energy dissipation device.

Point b in Fig. 5, associated with the onset of yielding in the energy dissipation device, is de-scribed by the lateral force Hyand lateral displacement ygiven by,

( )( )

++=

w

rrypedy

l

aHKKH 13.

(6)

and

( )

+

=

w

r

w

eff

ed

s

y

yl

a

al

hL

E

f132

(7)

Where fy is the yield strength of the steel in the energy dissipation devices.By substituting Eqs 3, 5 and 7 and simplifying, Eq. 8 is obtained:

+

+

= wrsdysd

sp

sp

ed

s

sp

eff

wy

l

aHAf

A

A

L

L

E

E

h

alH 131

2

1

(8)

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Fig. 6 Mechanical model for determining the response of a hybrid wall [14]

b c

a

fs u

Strain

Stress

fy

y

s u

2

s u

f -f

-

2

s u y

s u y

Es

1

1 E =s

O

Fig. 7 Idealised Stress-strain relationship for mild steel in tension [14]

Point c in Fig. 5, corresponds to the development of a stress equal to the ultimate tensile strength of

the energy dissipators in the idealised stress-strain relationship shown as point b in Fig. 7. The force H p

and displacement udescribing this point are given by the incremental equations,

yyppedp HKKH ++= .' (9)

and


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( ) yw

eff

ed

s

ysu

pal

hL

E

ff+

=

'2

(10)

Point c in Fig. 5marks the ultimate deformation capacity of the wall for use in design when incorporat-

ing axial energy dissipation devices. The displacement capacity beyond this point should not be relied upon

since the energy dissipators themselves will not be subjected to axial strains only but also to bending due tothe kinematics of the system [14].

The tangential stiffness between points cand din Fig. 5 equals K p. This means that the lateral force in

this region of the curve is given by,

ppp HKH += (11)

The self-centering response achieved in jointed systems enables hysteretic damping to be treated as if

viscous. That is, the system is ideal for use in a displacement-based design approach such as that pro-

posed by Priestley [3]. This approach takes a non-linear multi-degree-of-freedom (mdof) system and con-

denses it into a single-degree-of-freedom (sdof) elastic structure with a corresponding equivalent viscous

damping ratio. However, the level of equivalent viscous damping in jointed system cannot be uniquely de-

fined since it is amplitude dependent.

Consider the case of a wall prestressed with tendons and having axial energy dissipation devices at ornear the wall centerline. The energy dissipated, ED, by the devices in a complete cycle is from Fig. 8:

wdbuBdduBD lFFE == 24 (12)

where Fdis the force in the energy dissipation devices, d= b lw / 2 is the wall uplift at the centerline;factor u is the ratio between the energy dissipated by an ideal bi-linear dissipator and a rigid plasticdissipator with a force capacity Fd. Values of u range from 0 at the onset of yielding and tend asymp-totically to 1; Bis the ratio of the actual energy dissipated to that dissipated by the idealized bi-lineardevice. Thus, this factor accounts for the Bauschinger effect. Typical values of B for low carbonsteel range between 0.8 and 0.9.

The strain energy, ESo, in an elastic sdof structure is given by,

2=HESo

(13)

The energy dissipation devices begin to dissipate energy once they exceed the elastic limit. Thisimplies the opening of a gap at the base of the wall. At this stage the lateral displacement at theeffective height of the wall can be related to the base rotation, b:

effbf h= (14)

where is fa factor that considers all sources of flexibility in all elastic regions of the system: flexuraland shear deformations in the wall panel, in the foundation element and the flexibility in the systeminduced by the rotation of the foundation due to the soil-structure interaction. Factor f is largest at theonset of yielding of the energy dissipation devices. This factor tends to 1 as the opening of the joint

becomes the largest contributor to the imposed lateral displacement in the wall.Now, the equivalent viscous damping due to hysteresis, h, can be obtained as

So

Dh

E

E

4= (15)

Substituting Eqs. 12 to 14 into Eq. 15 and simplifying:

H

F

h

l d

eff

w

f

uBh

= (16)

but the lateral force H in the substitute structure can be approximated by:

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( )PNFh

lH d

eff

w ++=2

(17)

Substituting Eq. 17 in Eq. 16 and simplifying:

( ){ }dfuB

hFPN /1

12

++

=

(18)

The lower bound for h is obtained when u= 0 h,min = 0. This situation arises when the dissi-pators respond elastically. The upper bound for h is obtained from Eq. 18 by replacing u =1, B =0.9, f =1 and Fd= F du. Therefore,

( ){ }duh

FPN /1

573.0max, ++

= (19)

The requirement for self-centering response implies that the gap at the base joint of the wallscloses upon unloading. Thus, the prestressing force after losses, in addition to the gravity loads, mustbe sufficient to bring the energy dissipation devices to nominally zero strain upon unloading. Underreversed cyclic loading in the tensile strain domain, the compressive stress at zero strain is very simi-lar to the tensile stress obtained at the maximum reversal point for low carbon steels, see Fig. 8. Thislimiting condition implies that:

PNFdu + (20)

Fig. 8 Energy dissipated per cycle by a centrally located axial dissipator


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When Fdu = N+P, the maximum equivalent hysteretic damping is from Eq. 19, h,max = 0.287. Inpractice, however, the maximum values obtained for h,max are of seldom larger than 0.20.

The total equivalent viscous damping, eq, for design should also consider the intrinsic and radia-tion damping. These two sources of damping can be approximately taken as 0.05. These twosources contribute significantly to reducing the dynamic response at low amplitudes and less at largeamplitudes. Hence, the total equivalent damping in a hybrid wall is,

05.0+= heq (21)

5 DYNAMIC RESPONSE

The static response characterisation and subsequent use of the response in a displacement-based design approach with a sdof structure falls short of describing the dynamic response of a multi-degree of freedom system. Following a capacity design approach, it should be ensured that in jointedsystems the sole source of non-linear deformations are those due to the opening of the joints. Allother regions and elements in the structure should be made stronger to ensure elastic response. Forthis reason, it is necessary in design to complement the understanding obtained from the static re-sponse characterisation with the trends observed from the dynamic response. In particular, the bend-ing moment and shear force envelopes in the walls should be expected to be quite different from thoseestimated with any static procedure.

Fig. 9 shows the bending moment and shear force envelopes obtained from non-linear time his-tory analyses for a twelve-storey cantilever wall building subjected to a synthetic record matching adesign response spectrum. Analyses were conducted on models representing conventional monolithiccast-in-place reinforced concrete wall construction and the hybrid cantilever wall incorporating axialenergy dissipators described in Section 4.

To enable a comparison of the dynamic response, the backbone moment-curvature responseemployed in the analyses was identical for the monolithic and jointed wall models. The response ofthe monolithic wall was modelled using the Takeda hysteresis rule, whereas an Origin-centered rulewas used to represent the response of the hybrid wall. The monolithic wall was designed for ductileresponse following the recommendations of the New Zealand Loadings and Concrete Structures

Standards [15, 16]. The hybrid wall was designed to match the capacity of the monolithic wall. Theaxial load level for the walls was small and was ignored in the analyses. Constant 5% damping wasassigned to all modes of response.

Fig. 9 (a) shows that, for the particular case studied, the maximum bending moment demand in

(b) Shear force envelope(a) Bending moment envelope

0

1

2

3

4

5

6

7

8

9

10

11

12

0 2000 4000 6000 8000 10000 12000 14000 16000

Bending moment, kNm

Storey

Monolithic

Jointed

Mn = 11870

1

2

3

4

5

6

7

8

9

10

11

12

0 500 1000 1500 2000

Shear force, kN

Level

Monolithic

Jointed

NZS 3101:1995, . o = 1.46

Fig. 9Comparison between the dynamic response of monolithic and jointed walls in a twelve-storey building

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the hybrid wall develops above the base where the joint opens. In the case of the monolithic wall, themaximum bending moment develops at the base. Yielding in this wall spreads upwards to about 50%of the walls height, a value greater than currently assumed in design. In the upper half of the wallsthe bending moment envelopes are nearly identical for both systems.

Fig. 9 (b) compares the shear force envelopes obtained from the analysis of both systems withthe design envelope derived from the recommendations of the New Zealand Concrete Structures

Standard. The shear force envelopes are very similar for both systems and are, for most of the heightof the walls, are within the design envelope obtained from the standard. This finding suggests that theshear force magnification factor, recommended for the design of monolithic walls in this particularstandard, might also be equally applicable to hybrid walls incorporating energy dissipators.

6 EXPERIMENTAL WORK

Rahman and Restrepo [12], Holden et al. [14, 17] and Toranzo et al. [13] (also Toranzo [18]) havereported testing a variety of jointed wall systems.

Rahman and Restrepo tested three half-scale precast concrete wall units, representative of afour-storey building. The units were designed to ensure that yielding of the tendons would not takeplace below 2.5% drift. The wall specimens were 4 m high by 1.35 m long by 125 mm thick. Units 2

and 3 incorporated energy dissipation devices, whereas Unit 1 was post-tensioned only. Additionalgravity loading was simulated in Unit 3 by externally applying 200 kN concentric axial force through aservo-controlled load system. Fig. 10 shows the geometry of the test specimens and the loading

Fig. 10 Test set-up used in the testing of units by Rahman and Restrepo [12]


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frame. Quasi-static reversed cyclic loading wasapplied to the specimens by a horizontal actuatorfitted at the top of the walls. The specified concretecompressive strength was 45 MPa. The wall speci-mens were prestressed with 2-12.7 mm diameterlow relaxation strands. The strands were fed

through 70 x 20 mm flat ducts. The ample gap pro-vided by the ducts was deliberately chosen to mini-mise friction. The total prestressing force in theunits at the commencement of the tests was 188 kN.

The lowermost corners of the wall units wereconfined with an arrangement of closely spacedGrade 400 rectangular hoops and cross ties havinga nominal volumetric steel ratio = 2%, see Fig. 11.The confining cage was extended up to a heightequal to twice the neutral axis depth associated withan extreme compressive strain of 0.4%.

A 20 mm diameter Grade 430 steel bar with a

nominal strength of 430 MPa was used for the axialenergy dissipation devices in Units 2 and 3. Thediameter of the milled portion of the bar for Unit 3was 16 mm. The milled portion was 200 mm long.The milled portion was cast into the foundationbeam to ensure that the surrounding concrete pre-vent buckling from occurring. Fig. 12 shows theenergy dissipation devices placed for casting in thefoundation beam of Unit 3.

The lateral force-deflection response of Unit 1is shown in Fig. 13 (a). The non-linear elastic re-sponse is clearly evident in this figure. As expectedfor this system, no residual displacements were ob-

served to develop in the wall during the test. Theinitial stiffness of the wall specimen at small dis-

Fig. 11View of cage of Unit 3 [12]

Fig. 12Close-up view of energy dissipation devices [12]

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(a) Unit 1

(b) Unit 3

Fig. 13 Hysteretic response of a jointed and a hybrid wall tested by Rahman and Restrepo [12]

Fig. 14 Extent of damage in Unit 3 at 4% drift [12]


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placements was maintained even after excursions to large lateral drifts. In addition, no degradation inlateral strength capacity occurred with the progressively increasing lateral drift demands.

Fig. 13 (b) shows the hysteretic response of Unit 3. The hysteretic energy dissipation provided bythe axial energy dissipators is evident in the lateral force-lateral drift response. No strength degrada-tion occurred during testing below 3% drift. The prestressing strands, together with the simulatedgravity loading, provided a restoring force so the wall always returned after unloading to its original

position. After the tendons exceeded the limit of proportionality, some residual drift was observed. Theapparent loss of stiffness observed evident at large drift levels in Fig. 13 (b) was mainly due to irrecov-erable compressive strains developed in the confined concrete at the wall toes.

The extent of concrete cracking and spalling in the wall panel of Unit 3 near the end of the test isillustrated in Fig. 14. Cosmetic spalling of the concrete took place over a length of approximately 180

Fig. 16 Hysteretic response of the east energy dissipator in Unit 3 [12]

Fig.15 Equivalent viscous damping obtained from the hysteretic response of Unit 3 [12]

Unit 3

0

2

4

6

8

10

12

14

16

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5

Drift, h (%)

EquivalentViscous

Damping(%)

1st Cycle

2nd Cycle

3rd Cycle

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over 5.5 m. The elastic modulus for the tendons was Esp = 160 GPa. These tendons wereprestressed to 46% of the ultimate tensile strength to induce an initial prestressing force of 132 kN.

The concrete of wall panel in the hybrid unit included 35 mm long high strength steel fibres at adose of 40 kg/m

3. This was done to reduce conventional reinforcement in the wall panel. Axial energy

dissipation devices, similar to those used by Rahman and Restrepo but with 15.8 mm in diameter,were provided in this unit.

An arrangement of diagonal steel reinforcement welded to steel plates was designed to transferthe internal forces in the wall panel directly to the foundation, without the need for confining the con-crete at the wall toes, see Fig. 17.

(a) Monolithic wall unit at 2.5 drift (b) Hybrid wall unit at 3 drift

Fig. 18 View of wall panel reinforcement at large drifts [14, 17]

-180

-140

-100

-60

-20

20

60

100

140

180

-120 -90 -60 -30 0 30 60 90 120

Lateral displacement (mm)

Appliedlateralforce(kN)

Monotonic response

-3.0 -2.0 -1.0 0 +1.0 +2.0 +3.0

Drift (%)

Hn = 154 kN

Hn = 154 kN

(a)

(a) Monolithic unit

Fig. 19 Hystertic response of the units tested by Holden et al. [14, 17]

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The conventional precast reinforced wall performed very well in terms of ductility capacity and energyabsorption capability, reaching 2.5% drift before a number of fractures in the reinforcement resulted insignificant strength degradation. This unit, however, showed a significant number of residual cracks inthe cycles past 1% drift.

The hybrid wall unit achieved drift levels well in excess of 3% with no visible damage to the wallpanel prior to failure. Fracture of the energy dissipators was observed at 5% drift. The carbon fibretendons fractured at 6.2% drift. This suggests that this system can be designed to have large reserveof lateral deformation capacity.

Fig. 18 shows a close-up view of the wall panels at the base of the wall. It is evident that the de-velopment of plasticity in the monolithic wall is associated with extensive damage, whereas at largelevels of drift no visible damage was observed in the hybrid wall.

-120

-80

-40

0

40

80

120

-120 -90 -60 -30 0 30 60 90 120

Lateral displacement (mm)

Appliedlateralforce(kN)

-3.0 -2.0 -1.0 0 +1.0 +2.0 +3.0

Drift (%)

Theoretical pushover response

(b)

(b) Hybrid unit

Fig. 19Hystertic response of the units tested by Holden et al. [14, 17]


56


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Fig. 19 shows the lateral force-lateral displacement response of the units tested by Holden et al.The monolithic unit, Unit1, showed the characteristic fat loops whereas the hybrid unit, Unit 2,showed a characteristic self-centering response. Fig. 20 plots the residual drifts for the units. Themonolithic unit displayed significant residual displacements. These displacements were equal to thepeak displacement less the small elastic recovery. In contrast, the hybrid wall unit showed essentiallyno residual displacements.

In many countries the adoption of seismic design approaches combined with traditional methodsof construction has resulted in structural conflict. For example, very often the design of frame build-ings aims at developing the beam sidesway mechanism shown in Fig. 1. In actual construction, how-ever, the frames are infilled with unreinforced brick masonry. The use of unreinforced masonry infills

in the frames leads to a dramatic change in the seismic response. The presence of the infills pre-cludes the formation of the mechanism postulated in design. In turn, the brittle failure of an infill oftenresults in the formation of an intermediate column sideway mechanism and leads to extremely poorbuilding performance and even collapse.

Toranzo et al. [13] (also Toranzo [18]) reported the first shake table testing of a 40% scalethree-storey hybrid wall unit. The wall unit in this test was built with brick masonry infill, and includedthe slabs, see Fig. 21. The columns were designed to for strain control to ensure small shear distor-tions would occur in the wall panel, while the wall rocked at the foundation. Energy dissipation devices,in the way of tapered levers designed to yield in bending with constant curvature, were installed at thewall toes, see Fig. 22.

This test unit was subjected to 60 dynamic tests on the 200 KN capacity shake table at theUniversity of Canterbury, New Zealand. Fig. 23 shows the response of the unit under the Taft N21E

record of the 1952 Kern County earthquake record magnified to attain a target peak ground accelera-tion of 0.5g. The seismic response of this unit was excellent. The masonry infill maintained its integ-rity throughout. No cracking was observed to occur in the infill. The maximum residual drift of 0.13%was observed after an excursion to 1.8% drift. The residual drifts were due to the spreading of yieldlines in the slabs.

-2.5

-1.5

-0.5

0.5

1.5

2.5

-3 -2 -1 0 1 2 3

Peak drift (%)

Residualdrift(%

)

1st Cycle

2nd Cycle

3rd Cycle

Unit 1

Unit 2

(c)

Fig. 20Residual drifts measured during testing reported by Holden et al. [14, 17]

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shake table

4585

3980

1260

1260

126

0

200

130 580 130

120 470 840 470 120

Fig. 21 Details of the shake table test unit by Toranzo [18]

foundation beam

rocking wall

dissipator-end connectedto the wall throught pin

dissipator-end fixed to

foundation through bracket

brackets fixed

to foundation

Fig. 22 Details of the flexural energy dissipation devices [18]


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7 CONCLUSIONS

1. This paper described the behavioural concepts pertinent to a new generation of structural con-crete jointed systems.

2. The main advantages of the jointed systems described are the large lateral displacement capacity,the lack of structural damage associated with large displacements and the ability to return to the

original position upon unloading.

3. Jointed systems can be explicitly designed with hysteretic energy dissipation devices to provideenergy dissipation with a theoretical equivalent viscous damping up to 0.287.

4. The self-centering response characteristics of jointed systems make them ideal for use with adisplacement-based design approach. The equivalent viscous damping is, however, amplitudedependent. This dependence must be considered in design.

5. Results from non-linear time history analyses indicate that the dynamic response of hybrid wallsystems, as reflected by the bending moment and shear force envelopes, is similar to that ofconventional monolithic wall systems.

6. Quasi-static and shake table experimental work have clearly shown the benefit of jointed systems.

ACKNOWLEDGMENTS

The experimental part presented in this paper was carried out while the author on the faculty atthe University of Canterbury, New Zealand. The author is grateful to the colleagues and techniciansfor the support that made the experimental program successful.

Special recognition is also due to Mr. L. Toranzo, Ph.D. candidate and to Mr. T. Holden, ProjectEngineer, KPFF, San Francisco, who as students at the University of Canterbury conducted part of theresearch and provided material for this paper.

Finally, the author expresses his gratitude to Prof. A. Rutenberg, Technion University, Israel, forproof reading this manuscript.

Run 41

(Taft x 2.83)

-40

-20

0

20

40

60

-1.5 -1 -0.5 0 0.5 1 1.5 2

Drift (%)

Basemoment

(kNm

)

Fig. 23 Base moment vs drift dynamic response to the magnified Taft earthquake record [18]

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