High-Force-to-Volume Seismic Dissipators Embedded in aJointed Precast Concrete Frame

Geoffrey W. Rodgers1; Kevin M. Solberg2; John B. Mander3; J. Geoffrey Chase4;Brendon A. Bradley5; and Rajesh P. Dhakal, M.ASCE6

Abstract: An experimental and computational study of an 80-percent scale precast concrete 3D beam-column joint subassembly designedwith high force-to-volume (HF2V) dampers and damage-protected rocking connections is presented. A prestress system is implementedusing high-alloy high-strength unbonded thread-bars through the beams and columns. The thread-bars are posttensioned and supplementalenergy dissipation is provided by internally mounted lead-extrusion dampers. A multilevel seismic performance assessment (MSPA) is con-ducted considering three performance objectives related to occupant protection and collapse prevention. First, bidirectional quasi-static cyclictests characterise the specimen’s performance. Results are used in a 3D nonlinear incremental dynamic analysis (IDA), to select criticalearthquakes for further bidirectional experimental tests. Thus, quasi-earthquake displacement tests are performed by using the computation-ally predicted seismic demands corresponding to these ground motions. Resulting damage to the specimen is negligible, and the specimensatisfies all performance objectives related to serviceability, life-safety, and collapse prevention. DOI: 10.1061/(ASCE)ST.1943-541X.0000329. © 2012 American Society of Civil Engineers.

CE Database subject headings: Damping; Seismic effects; Damage; Earthquakes; Dynamic analysis; Precast Concrete; Frames.

Author keywords: High force-to-volume damper; HF2V; Lead extrusion damper; Multi-level seismic performance assessment; Damageavoidance design; Quasi-earthquake displacement testing; Incremental dynamic analysis.

Introduction

Research and development of jointed precast concrete structureshas gained considerable momentum, with significant research onthe PRESSS systems (Priestley et al. 1999; Stanton et al. 1997)and Damage-Avoidance Design (DAD) systems (Ajrab et al.2004; Holden et al. 2003; Mander and Cheng 1997). These systemsaccommodate inelastic behavior by rocking at specially detailedjoints; provide a level of seismic resistance comparable to currentstandards for ductile (damage-prone) structures; and remain essen-tially damage-free without the excessive residual displacementcommon in conventional systems. Residual displacement is critical

because it can lead to complete postearthquake loss of structuralamenity. However, these systems may exhibit relatively low hys-teretic energy dissipation resulting in excessive response motion.Although this may not be an issue with long period structures,short to medium period rocking structures may undergo signifi-cantly greater displacements than conventional ductile structures(Priestley and Tao 1993). Hence, ductile jointed precast concretesystems ideally require some supplemental energy dissipation todissipate earthquake energy and mitigate an excessive displacementresponse.

Energy dissipation devices of varying sophistication are avail-able. Early applications in ductile jointed connections provided dis-sipation by mild steel reinforcing bars grouted in ducts across thejoint (Stanton et al. 1997). The devices proved to work well withthe grout providing a degree of buckling resistance. However, de-bonding of the bars under cyclic loading caused some stiffness deg-radation. Other options include tension-only mild steel devices(Bradley et al. 2008) and external mild steel devices allowed tobuckle (Li et al. 2008; Solberg et al. 2008).

Each option has limitations. Mild steel devices allowed tobuckle lose significant energy dissipation upon buckling, givingmuch less dissipation on subsequent cycles. Buckling restraineddevices may provide more stable resistance, but considerableresidual compression forces remain upon joint closing. Residualforces from these elements will effect initial structural stiffnessand have the possibility of a negative impact on response. As such,residual forces from damping elements must be considered in de-sign. Mild steel devices can also suffer failure resulting from low-cycle fatigue and should strictly be replaced following a significantearthquake. Finally, all yielding-and buckling-based dissipatorsprovide reduced capacity on subsequent smaller response cycleswhere no yielding may occur.

Other research has looked at alternatives for providing energydissipation to concrete structures. Pekcan et al. (1995) examined

1Postdoctoral Researcher, Dept. of Mechanical Engineering, Univ. ofCanterbury, Private Bag 4800, Christchurch 8140, New Zealand (corre-sponding author). E-mail: geoff.rodgers@canterbury.ac.nz

2Master’s Student, Dept. of Civil Engineering, Univ. of Canterbury,Private Bag 4800, Christchurch 8140, New Zealand.

3Zachry Professor of Design and Construction Integration, Zachry Dept.of Civil Engineering, Texas A&M Univ., College Station, TX 77843.E-mail: jmander@civil.tamu.edu

4Professor, Dept. of Mechanical Engineering, Univ. of Canterbury, Pri-vate Bag 4800, Christchurch 8140, New Zealand. E-mail: geoff.chase@canterbury.ac.nz

5Lecturer, Dept. of Civil Engineering, Univ. of Canterbury, Private Bag4800, Christchurch 8140, New Zealand. E-mail: brendon.bradley@canterbury.ac.nz

6Associate Professor, Dept. of Civil Engineering, Univ. of Canterbury,Private Bag 4800, Christchurch 8140, New Zealand. E-mail: rajesh.dhakal@canterbury.ac.nz

Note. This manuscript was submitted on April 8, 2008; approved onOctober 6, 2010; published online on October 21, 2010. Discussion periodopen until August 1, 2012; separate discussions must be submitted forindividual papers. This paper is part of the Journal of Structural Engineer-ing, Vol. 138, No. 3, March 1, 2012. ©ASCE, ISSN 0733-9445/2012/3-375–386/$25.00.

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the use of elastomeric spring dampers to provide energy dissipa-tion. Christopoulos et al. (2008) developed a self-centering, energydissipating bracing system that could be used in either concrete orsteel structures; whereas, Shen et al. (1995) examine the use of vis-coelastic dampers to provide seismic resistance to concrete frames.The development of passive control systems that use yielding steelbraces and shape-memory alloys is detailed in Dolce et al. (2005).

High force-to-volume (HF2V) lead-based damping devices pro-vide an attractive alternative and consist of a central shaft with astreamlined bulge encased in lead. When the shaft moves, the leadyields because of the drag effect created by the bulge (Rodgers et al.2007). Properly designed, these devices behave like a type of Cou-lomb damper with a slight velocity sensitive effect. Unlike mildsteel devices, this velocity sensitivity attribute permits the HF2Vdevice to creep back to near zero force over time providing a reus-able device with no maintenance required following an earthquake.More importantly, HF2V devices can provide the same resistivecapacity every response cycle.

Historically, lead-based damping devices were quite large limit-ing applications to locations like base isolation (Cousins and Porritt1993). Recent research has developed volumetrically small deviceswith the same force capacity capable of direct placement in beam-column joints (Rodgers et al. 2007; Rodgers et al. 2008b). As thesedevices have some unique manufacture and assembly attributes,they are referred to in this paper as a high force-to-volume(HF2V) damper.

On the basis of lead-extrusion damper (Cousins and Porritt1993; Robinson and Greenbank 1976) and HF2V damper (Rodgerset al. 2008a) tests, the velocity exponent, α≈ 0:11–0:13, where:

F ¼ Cαvα ð1Þwhere F = the damper force; v = the velocity of the shaft; Cα = aconstant dependent on the device architecture; and α = the velocityexponent. More detailed velocity dependence studies can be foundin Mander et al. (2009), and Rodgers (2009).

This research incorporates the findings from previous studies(Bradley et al. 2008; Li et al. 2008; Rodgers et al. 2007; Rodgerset al. 2008b) and focuses on the development of cost effective reli-able energy dissipation and detailing schemes. Previous work re-lated to external dissipation devices. This research places the HF2Vdampers directly inside the concrete beam-end-zone providingreliable energy dissipation and an architecturally pleasing finish.Detailing of the beam-ends that frame into the column is modifiedto accommodate the devices. A dual experimental-computationalstudy investigates the seismic capacity of the proposed jointed pre-cast concrete frame system versus the seismic demands expectedfrom a variety of adverse earthquake scenarios.

Subassembly Development

Construction Considerations

Special attention was given to potential construction issues. Thebeam and column elements were designed to be precast with lim-ited concrete placement required on-site. A cast in situ ‘closurepour’ was provided at one end of each beam. The other endwas considered to be a “dry-joint.” Hence: (1) the closure pour pro-vides an access point for coupling the prestress thread-bars and thedamping devices; and (2) high performance concrete can be used inthe high stress zone at the beam-end. The closure pour, thus, be-comes the primary focus for on-site erection. Within this region, theHF2V damper in each beam can be coupled to a threaded rod anch-ored in the column, the prestress thread-bars would be coupled to

one another, and the armoring channels would be tightened againstthe face of the column. The damper within the beam at the dry-jointend can be located and cast within the beam off-site. The dampershaft connection can then be made from the opposing side of thecolumn face. A detailed overview of the construction sequence isgiven in Solberg (2007).

Specimen Design

A 3D subassembly representing an interior joint on a lower floor ofa ten-storey building was developed. The subassembly consisted oftwo beams cut at their midpoints and an orthogonal beam cut at itsmidpoint, which is the approximate location of the point of contra-flexure for seismic loads. These beams were all connected to a cen-tral column. The orthogonal beam, referred to as the gravity beam,was designed to support one-way precast flooring. The other twoseismic beams were designed for predominantly seismic forces. Itis important to note that the specific tendon drape utilized for thegravity beam must be modified to avoid having an end eccentricitythat cannot be permitted. The dimensions of the prototype memberswere 875 mm square columns, 700 × 500 mm beams, and a 3.6-m-storey height. The prototype moment capacity of the beam-end atthe column face was 500 kNm. The subassembly was scaled to80% of the prototype framed structure. The column was scaledto 700 mm square and the beams scaled to 560 by 400 mm.Fig. 1(a) shows a schematic of the building from which the speci-men was derived. Fig. 1(b) shows the experimental setup, wherefurther details on the specimen dimensions and design detailsmay be found in Solberg (2007), Bradley et al. (2008), and Li(2006). Fig. 1(c) illustrates the reinforcing layout for the structuralmembers, in which the target longitudinal reinforcement ratio is0.01. Note that in Fig. 1(c) both the seismic beams are shown.These beams are identical but both are shown to provide additionalsection views. These members were designed to remain elasticunder the expected rocking connection strength from the HF2V de-vice and prestress forces. Four D20 (f y ¼ 500 MPa) rolledthreaded, low-strength reinforcing bars (Reidbar™) are used forlongitudinal reinforcing at the top and bottom of the beam to pro-vide a moment capacity of ϕMn ¼ 260 kNm. Because of the axialprestress load, minimal transverse steel requirements governed.Thus, HR12 (f y ¼ 500 MPa) stirrups were provided in the beamat a spacing of half the beam depth (250 mm) and a closer100-mm spacing at the ends. Additional transverse reinforcementwas provided top and bottom 1.2 m from the beam-ends to confinethe concrete in these high compression (burstingstress) zones.

Fig. 1(c) also illustrates the column reinforcing layout. Alongitudinal reinforcement ratio of 0.01 was provided by 12D20 (f y ¼ 500 MPa) rebars and transverse steel consisting ofHR12 (f y ¼ 500 MPa) stirrups at a 250-mm spacing. The stirrupswere doubled and the spacing was halved within the joint region.Shear resistance was primarily provided by the core concrete attrib-uted to the axial load in the column. The stirrups were designedconsidering the expected overstrength of the jointed/rockingconnection.

Two 45-mm longitudinal PVC ducts spaced 200 mm apart wereprovided for the prestress system at the seismic beam vertical cen-trelines. Two 26.5-mm (MacAlloy, f y ¼ 1;100 MPa) thread-barsprovided prestress in the seismic direction utilizing a straight pro-file to ease congestion in the column and provide a more construct-ible solution. The thread-bars in the gravity beam were draped toprovide “load-balancing” for the gravity loading from the one-wayfloor panels; hence, this thread-bar crosses the joint’s centreline at a30-mm eccentricity.

A 300-mm cast in situ ‘wet’ joint was provided at the columnconnection end of each beam. The detailing strategy in the seismic

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direction is shown in Fig. 2(a). The joint was designed to accom-modate 150 × 150 mm HF2V devices at midheight in thebeam-end-zone in the seismic beams and at a 50-mm vertical offsetfrom midheight in the gravity beam. This offset prevents interfer-ence between the damper anchor rods from the two orthogonal di-rections in the column. The dampers are intended to be included inall beam-end-zones that connect to exterior columns. A 180-mmparallel flange channel was used top and bottom of the beamsto armor the contact surfaces and prevent crushing of the concreteat the rocking edge. The channel was provided to carry the concen-trated load at the rocking edge and mechanically distribute thisforce through the reinforcing bars into the beam. This armoring

concept prevents the large compression forces at the rocking edgecausing localized crushing of the concrete.

The channels also serve as a means of mechanically developingthe longitudinal reinforcing, by providing cuts on the interiorflange, whereby, the threaded longitudinal steel could be lockedinto the channel by using nuts. These nuts also provided a meansof ensuring the channel is flush with the column face during on-sitefabrication. Four 25 × 10 × 500-mm rods were welded in the cor-ners of each flange to help stiffen the joint region and ensure rigidrocking behavior. Finally, four 1-m threaded rods were spaced at100-mm centers to provide attachment and anchoring points for theHF2V devices.

2@10

0

1625

3200

280

4@10

03@

250

2x4@

100

HR

12 @

250

560

1015

700COLUMN

EAST SEISMIC BEAM

3850

CENTERLINE ROD TO BOTTOM OF BEAM AT MIDPOINT:

180mm @ RADIUS = 16.5m

REIDBAR COUPLER NUT, TYP

180x75x20.9 STEELCHANNEL, TYP

(2) 26.5mm HIGHSTRENGTH THREADEDROD IN 45mm PVC, TYP

4@100 4@200

4@2004@100

300

560

001@5052@21RH

GRAVITY BEAM

WEST SEISMIC BEAM

5@100HR12 @250

3850

HR12 @250

(4) D20 REIDBARLONGITUDINAL, TYP

5@100

300

4@200 4@100

3850300

A

AB

BCAST INSITU WET JOINT

ANCHOR RODS FORLE DAMPER, 18mm

A-A

B-B

560

400

SECTIONAL VIEW A-A

SECTIONAL VIEW B-B

230

400

560

250

100

END ELEVATION

CENTERLINE OF PRESTRESSING THREADBAR TO BOTTOM OF

BEAM AT MIDPOINT:180mm @ radius = 16.5m

CAST INSITU WET JOINT

(4) 18mm DIAMETER THREADED RODS TO

LOCATE HF2V DAMPER

3850300

4@100 4@200 HR12 @250 5@100

GRAVITY BEAM

WEST SEISMIC BEAM

(2) 26.5mm HIGH STRENGTHPRESTRESSING THREAD-

BARS IN 45mm PVC, TYP

REIDBAR COUPLER NUT, TYP

180x75x20.9 STEEL CHANNEL, TYP

(4) D20 LOW-STRENGTHROLLED THREADED

LONGITUDINAL REINFORCING BARS

(REIDBARSTM), TYP

4@100

300 3850

EAST SEISMIC BEAM3850 300

HR12 @2505@100

4@200 HR12 @250

B

B A

A

4@200 4@100

SECTIONAL VIEW B-B

SECTIONAL VIEW A-A

400

5@100

560

560

400

230

250

2@10

03@

250

2x4@

100

HR

12@

250

4@10

0

28056

016

25

3200

1015

100

700COLUMN

INCLINED POST-TENSIONING DUCTS

HORIZONTAL POST-TENSIONING DUCTS

N

S E

W

9

(a) (b)

(c)

6009600

96000

01 00

00

1 00

00

1 00

(SEISMIC DIRECTION)

(GRAVITY DIRECTION)

Fig. 1. Subassembly development and reinforcing detail: (a) prototype structure showing the location of subassembly [Reprinted with permissionfrom Solberg (2007)]; (b) photograph of the subassembly in experimental test apparatus; inset: Specimen Orientation (image by authors); (c) re-inforcing details of the beams and column; Note that the column has horizontal posttensioning ducts in one direction and inclined ducts in theorthogonal direction resulting from draped tendon profile in the gravity beam

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Shear loads were carried by four 30-mm shear keys located ateach corner of the connecting beam. In the beam, two 30-mm holeswere drilled in the top and bottom flange to provide a femaleattachment for the shear keys. The shear keys were tapered inward5° to prevent binding with the beam during connection opening. Inthe column, these shear keys were designed to be screwed into a nutlocated behind a hole in the column’s steel plates.

One of the primary objectives of this study was to improvebeam-column joint detailing by improving constructability andreducing materials compared to past designs (Li et al. 2008).Fig. 2(b) shows that column detailing was improved in that:(1) the column contact plates were reduced from a single full-depth-plate to two end-plates; (2) the prestress ducts ran straightrather than angled, reducing congestion; and (3) the internal damp-ers were connected using a single rod with grouting tubes. The col-umn end-plates were sized to provide a full contact surface for thebeam armoring and provide a 10-mm extension on all sides. Thisplate was checked to ensure concrete crushing in the column didnot occur at the design strength of the connection, and it was de-veloped into the joint core by using weld studs.

Specimen Construction

The specimen was constructed in several parts. First, the two seis-mic beams were caged with a cast insitu end, as shown in Fig. 3(a).This task required providing stubs for the longitudinal steel, theprestress system ducts, and the threaded rods for the HF2V damper.Next, the gravity beam and column were caged and cast. The grav-ity beam required the draped profile, and the damper and tendonswere offset 30 and 50 mm from centreline, respectively, to avoidintersecting the orthogonal ducts in the column.

The column longitudinal reinforcement was welded to a 20-mmsteel plate at the top and bottom, and the armoring plates were de-veloped into the concrete by welds studs. A 50-mm recess at thecolumn face where the damper anchoring rods are bolted allowedeasy bolt clearance. Corrugated steel tubing provided the damperducts where the shaft and connecting rods passed through. Grouttubes are provided to grout the anchoring rod within the column,and the damper duct in the beam-ends was left open to allow freemovement of the damper shaft.

The remaining elements were installed to enable the closurepour, as shown in Fig. 3(b). Within the closure pour, the HF2Vdevices were attached to the threaded rods in the beam. The dampershaft was coupled to the threaded rod in the column that was anch-ored against a steel washer in the recess on the column face. The

damper shaft-, coupler-, and column- threaded rod were all encasedin a duct and waterproofed from the concrete.

Once the damper was in place, the main prestress thread-barswere coupled together. Next, the channel top and bottom was tight-ened against the column face and locked with longitudinal steelnuts. Finally, a thin sheet provided a barrier between the wet con-crete and the column face to prevent any bonding. Note that thismeasure may not be necessary because the tensile capacity of theconcrete would be negligible and was taken merely as a precaution.

A high performance concrete mix designed for good workabilityand high-strength was used for the three closure pours. Steel fibers(2% by weight) were incorporated to help impede crack propaga-tion. This concrete had a measured 28-day compressive strength of76 MPa with 50 MPa measured for the regular beam and columnconcrete.

HF2V Dampers

The HF2V damper designs are shown in Fig. 4(a). Damper shaftmotion plastically extrudes lead between the shaft and outer bodyto dissipate energy, as shown in Fig. 4(b) (Rodgers et al. 2007). Asingle HF2V damper is placed within each beam-end, as shown in

Fig. 2. Joint region detailing: (a) closure pour region at beam-end-zone; (b) column detailing at the joint region

Fig. 3. The cast in situ joint showing: (a) the east beam with the thread-bar exposed; (b) east beam with the thread-bar enclosed in PVC andshowing the damper (images by authors)

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Fig. 2(a). The damper shaft [Fig. 4(a)] was designed to be coupledto a threaded rod in the column of the same size. Four 18-mm(f y ¼ 300 MPa) threaded rods at 100-mm centers were cast intothe precast beam to anchor the device within the closure pour,through oversize attachment holes on the devices that allow adjust-ment when coupled to the threaded rod.

To ensure the system recentres, the expected negative momentcontribution from the HF2V device in compression should notexceed the expected positive moment contribution from the pre-stress. Hence, the moment contribution ratio, λ, must be greaterthan 1:

λ ¼ ϕMPS;i

ΩdissMdiss> 1 ð2Þ

whereMPS;i = moment contribution from the initial prestress force;Mdiss = moment contribution from the dissipation devices in com-pression; Ωdiss = overstrength factor of the dissipation devices (as-sumed as 1.5); ϕ = understrength factor for the prestress (assumedas 0.85). Rearranging terms gives a ratio of moment contribution

from prestress and HF2V devices. Initial posttensioning of 250 kNper thread-bar gives a total of 500 kN at the interface. The damperswere designed for a 250 kN yield force, corresponding to λ ¼ 1:23and 1.06 in the north-south direction for positive and negative mo-ment, respectively, and λ ¼ 1:13 in the east-west direction.

Fig. 4(c) presents experimental force-displacement responses ofthe HF2V devices in a universal (Avery) testing machine. Devices1, 2, and 3 were installed in the west, east, and south joints of thespecimen, respectively. The devices provide a similar response withan average yield force of 270 kN. Differences in initial stiffness andyield force can be attributed to small voids forming in the lead re-sulting from incomplete prestressing to remove microvoids, whereRodgers et al. (2007) discusses this issue in detail.

Predicted Response

The analytical prediction method assumes members remain elasticand that once gap-opening of the beam-to-column joints occurs,displacements can be determined from rigid-body kinematics. Tak-ing the internal moment arm (ePS, ediss) from the rocking edge andsumming each contribution yields

M ¼X

MPS þX

Mdiss ð3Þ

where MPS = moment in the thread-bars (¼ PPSePS); and Mdiss =moment in the dissipation device (¼ Pdissediss). The force in thethread-bars is then obtained from

PPS ¼ Pi þAPSEPS

LtePSθconn ð4Þ

where Pi = initial posttensioning force (250 kN); EPS = thread-barelastic modulus (170 GPa); APS = thread-bar cross sectional area(552 mm2); Lt = unbonded length of thread-bar (9 m); θcon = con-nection rotation; and n = number of joint openings spanned by thethread-bar (2 in the east-west direction and 1 in the north-southdirection). The force in the HF2V device is thus

Pdiss ¼ min jKdissedissθconPy;diss

ð5Þ

where Kdiss = stiffness of the dissipation device measured as200 kN=mm, which was the average of the three devices; andPy;diss = “yield force” of the HF2V device (250 kN).

A nonlinear moment-rotation relationship was developed for thesubassembly. The behavior of the structural elements was assumedto be elastic, but partial separation of the beam from the columnface causes geometric nonlinearity from rigid-body rotations.Given an elastic moment-rotation analysis, the force-displacementresponse of the subassembly can be evaluated.

In the east-west direction, the horizontal force at the top of thecolumn, V col, can be found as the following given the moment atthe joint:

V col ¼ 2ML

LbLcð6Þ

where L = beam length to column centreline (9.8 m); Lb = clearsupport length of the beam (9.1 m); and Lc = storey height(2.8 m). The total top displacement of the system, given V col,can be attributed to localized joint rotation and the total elastic de-formation of the system:

Δ ¼ Δelastic þ θconLbLLc ð7Þ

ø30

ø48 14

5

102

(a)

(b)

(c)

Lead Filled ChamberLead Flow

Shaft Motion

Endcap

-400

-300

-200

-100

0

100

200

300

400

-2 0 2 4 6 8 10 12 14 16 18

HF

2V D

amp

er F

orc

e (k

N)

Displacement (mm)

Damper 1Damper 2Damper 3

Fig. 4. HF2V Damper details: (a) exploded isometric view; (b) crosssectional view; (c) hysteresis loops

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in which Δelastic = the elastic deformation of the system from flex-ure, given by

Δelastic ¼V col;uplift

12

�ðLc � DÞ3EI�col

þ L2cL3bL2EI�bm

�ð8Þ

where EI�bm and EI�col = effective stiffness of the beam(0:25EIbm;gross) and column (0:6EIcol;gross), respectively; andD = the depth of the beams (560 mm). The values of effective stiff-ness for the beam and column (0:25EIbm;gross and 0:6EIcol;gross) arebased upon the analysis in Li (2006). During rocking, an eccentricpoint force is applied to the rocking edge, and there is a portion ofthe structural elements that do not contribute to the bending resis-tance. The values of the effective stiffness are derived on the basisof St. Venant’s principle in Li (2006) and used in Solberg et al.(2008, 2009). A more detailed derivation of Eq. (8) can be foundin either Li (2006) or Rodgers (2009).

Experimental Setup

The experimental test setup shown in Fig. 1(b) is closely similar tothat in Bradley et al. (2008). The column was pinned to the floorusing a universal joint. Additional pins were on stiff struts nearthe end of each beam. Actuators A and B were located at thetop of the east and south face of the column, inducing displace-ments in the seismic (east-west) and gravity (north-south)directions, respectively, as shown in Fig. 1(b). Actuator C wasorthogonal to the side face of the gravity beam in-line with the grav-ity beam support strut and was used primarily to stabilize the speci-men. Rotary potentiometers were installed against the opposite faceof each actuator. An additional actuator at midheight of the gravitybeam simulates the precast one-way floor panels with a constant120 kN load spread over a 1.5 m timber block.

At one end of each prestress thread-bar anchor, load cells mea-sure the forces in the thread-bars. Four 32-mm high-strengththread-bars along the longitudinal column axis were stressed to500 kN for a 2000 kN total axial load (0:1f 0cAg), simulating gravityforces. A total of 24 potentiometers measured localized displace-ments. Within the joint, each coupler connecting the HF2V shaft tothe threaded rod in the column was converted to a load cell con-sisting of eight strain gauges compensating for bending and tem-perature effects. Four strain gauges were placed on the top andbottom web of the beam armoring channels to detect any potentialyielding upon gap-opening.

Test Methods

Displacement-controlled uni- and bidirectional testing was per-formed. Both quasi-static testing using cyclic loading patternsand quasi-earthquake displacement (QED) tests (Dutta et al.1999) using load patterns from computational simulation of the full10-storey prototype structure were employed. The QED method isintended to produce realistic displacement demands representativeof an expected seismic response (Dhakal et al. 2006).

Quasi-Static Displacement Profiles

Preliminary low drift level quasi-static tests were used to character-ize the specimen for use in a computational model for the QED test(Solberg 2007). Owing to the damage-free nature of the specimen,it was possible to conduct preliminary tests (both uni- and bidirec-tional) without damaging the specimen. Further quasi-static testingwas carried out following the QED testing regime, consisting ofuni-directional tests in each direction to a maximum drift of 3%

(equivalent to a radial drift of 4%) and bidirectional tests to a maxi-mum (radial) drift of 4%.

Quasi-Earthquake Displacement Testing

The QED method is intended as a more realistic protocol by cap-turing specimen behavior under ‘real’ earthquake ground motion.Unlike quasi-static testing, which uses controlled cyclic displace-ments in ascending order, QED testing captures small loadingcycles following severe displacement demand from initial pulses.Second, P�Δ effects can be considered in the analytical model,thus, capturing nonuniform displacements resulting from excessiveyielding in a single direction. Finally, QED test performance can beextrapolated to infer likely damage at multiple levels of excitation.

Quasi-static test data was used to create an equivalent computa-tional model of the specimen. Details of the development of the 3Danalytical model are given in Bradley et al. (2008). Elastoplasticand bilinear elastic springs represent the dampers and prestressat the joint, respectively. The natural period of the full structurewas found to be 1.5 s.

To perform a Multi-Level Seismic Performance Assessment(MSPA), the earthquake records to be used must be preidentified.Dhakal et al. (2006) proposed a methodology on the basis of In-cremental Dynamic Analysis (IDA), in which an IDA is conductedusing multiple earthquake ground motions and the IDA results areprobabilistically processed to select records that give medium andhigh confidence at desired levels of seismic intensity. Performingan IDA involves conducting nonlinear dynamic analyses of a com-putational structural model subjected to a suite of earthquakeground motion records scaled to different intensity measures(IMs) (Vamvatsikos and Cornell 2002). For each analysis, an en-gineering demand parameter (EDP) is monitored, producing anIDA curve (i.e., a plot of IM versus EDP) for each earthquakerecord.

A reliable computational model of the structure yields displace-ment profiles at nodes of interest for use in physical testing. Thistask required the identification of earthquakes likely to representvarious levels of demand considering both rare and relatively fre-quent earthquakes. The procedure described by Dhakal et al. (2006)was adopted to define three key earthquake records representingmultiple levels of seismic demand, by performing an IDA(Vamvatsikos and Cornell 2002) to identify the structural responsefrom various earthquakes. Earthquakes representing percentilelevels at various intensities can then be identified and used forsubsequent analysis.

Assuming a firm soil site in Wellington, New Zealand (a highseismic zone), three levels of demand were identified followingDhakal et al. (2006). These demand levels were: (1) a 90th percen-tile design basis earthquake (DBE); (2) a 50th percentile maximumconsidered earthquake (MCE); and (3) a 90th percentile MCE. TheDBE and MCE were defined as earthquakes with return periods of475 years (10% in 50 years) and 2475 years (2% in 50 years), re-spectively. For the site of interest, this corresponds to a peak groundacceleration (PGA) of approximately 0.4 g for the DBE and 0.8 gfor the MCE, on the basis of the seismic hazard model in Stirlinget al. (2002). For each demand level, a performance level related toserviceability and life-safety was defined. For the 90th percentileDBE, it corresponds to a high level of confidence that the structureremains operational. After the MCE the structure is expected to berepairable with a moderate level of confidence (50th percentileMCE) and is not expected to collapse with a high level of confi-dence (90th percentile MCE).

Earthquake records were selected from a suite of 20 recordedground motions from the SAC project (Somerville et al. 1997).The spectral acceleration (Sa) at the fundamental period of the

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structure was selected as the intensity measure (IM) (Baker andCornell 2005). Thus, the DBE andMCE intensity levels correspondto a Sa of 0.27 and 0.48 g, respectively. The resulting IDA data isplotted in Fig. 5, showing the 10th, 50th, and 90th percentilecurves, which notes the three selected records. These records cor-respond to peak (radial) interstory drifts of 1.6, 1.6 and 2.8% for the90% DBE, 50% MCE, and the 90% MCE, respectively.

Experimental Results

This section first presents results from the quasi-static tests fol-lowed by QED test results.

Quasi-Static Test Results

The response of the subassembly in the east-west and north-south direction is given in Fig. 6. The specimen was subjectedto two fully reversed displacement cycles, at column drift ampli-tudes of 0.5, 0.75, 1.5, 2.25, and 3.0%. The prediction by usingEqs. (3)–(8) is plotted along with the experimental results, withgood agreement between the two.

In the east-west direction, the hysteretic behavior displaysnotable energy dissipation. The dissipation is notably high, giventhat the structural members remain essentially elastic; therefore,providing essentially no hysteretic energy dissipation. Further-more, the straight tendon profile in the east-west (seismic) direction

results in very low friction between the duct and tendons; therefore,also contributing very little inherent energy dissipation.

The hysteretic results without damper contributions are pre-sented in Rodgers (2009) and show an essentially bilinear elasticresponse regime with negligible energy dissipation. Evaluation ofthe energy dissipation provided by HF2V dampers in a damage-avoidance steel connection has been presented in Mander et al.(2009). Rodgers et al. (2008b) presents a comparison of the energydissipation achieved with a connection similar to the one in thismanuscript, comparing the connection without supplementaldamping with HF2V dampers and with yielding steel fuse bars.

The specimen did not suffer any noticeable stiffness or strengthdegradation, and stable hysteretic energy dissipation is evident. Themaximum recorded residual drift was approximately 0.08%, (2.6%of the maximum drift) and is attributed to friction arising within theprestressing system.

Fig. 6 shows that the yield drift of the experimental specimen inthe east-west direction is approximately 0.5%. Although the beamsused in the experimental specimen have a relatively high slender-ness ratio, they are posttensioned, prestressed concrete, which af-fects their behavior. The yield drift of 0.5% is in-line with whatwould be expected for structures with notably lower slendernessratios. The experimental specimen was based upon a well-knownNew Zealand design (Bull and Brunsdon 1998) and for NewZealand seismicity levels, which are very similar to those inCalifornia.

One aspect of the results in Fig. 6 that should be noted is theamount of added lateral force (applied column-shear force) attrib-uted to the HF2V dampers. Fig. 4(c) presents the hysteresis loopsfor the HF2V damping devices showing that they produce peakforces of around 250–300 kN. However, Fig. 6 shows that the en-closed area within the flag-shaped hysteresis loops spans approx-imately 100 kN, which represents �50 kN from what the responsewould be without dampers. This observation is simply attributableto the small lever arm from the rocking edge present for the damp-ers, as they are located at midheight of the beam. Therefore, oncethe equivalent moments are calculated, 300 kN resistive force ap-plied at the line of action of the dampers corresponds to 60 kNlateral load at the top of the column. Details of connectionmodeling and specific geometric relationships can be found inRodgers (2009).

In the north-south direction, the results are similar with stableenergy dissipation. The maximum residual drift was 0.12%, (4% ofthe maximum). This value is slightly larger than that in the

Fig. 5. Incremental dynamic analysis curves resulting from theanalytical model

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east-west direction and is attributed to the increase in prestress fric-tion from the draped tendon profile. Differential friction forces re-sulting from reversed cyclic loading contribute to this effect.

These frictional effects are evident in Fig. 7(a), which shows thechange in prestress force for a given column displacement. It isevident that friction between the duct and prestress thread-barsleads to some energy dissipation that is not considered in the origi-nal design. This effect is minimized in the east-west direction wherethe straight ducts result in less tendon-duct friction. Fig. 7(b) showsthe in-service response of the HF2V damper in the east beam. Theapparent in-service stiffness of the damper was reduced comparedto the damper tests alone. It is very important to note that the dis-placement in Fig. 7(b) is not directly measured but is, rather, aninferred displacement measured across the joint at the beam mid-height by a potentiometer attached to the column face. It, therefore,records the midheight gap-opening displacement and includes allsources of flexibility from axial stretching of the connecting ele-ments to take-up in the threaded coupler nuts. Because of the em-bedded device design, direct in-service device displacements wereunavailable. These additional sources of flexibility and slack re-duced the contribution of the damper to the overall hysteretic per-formance on smaller cycles. For example, the HF2V device should“yield” at approximately 1 mm of elongation but was not observeduntil ∼2:5 mm. The revised stiffness of the HF2V devices consid-ering the added freedom in the connecting elements is ∼80 and50 kN=mm for the gravity and east-west joints, respectively.

Although improved detailing of the damper connecting ele-ments will reduce the effects of added flexibility, these effects willalways be present and can have a notable influence on the dampercontributions to the overall subassembly response. The sources offlexibility are not discussed here but are discussed in detail inRodgers et al. (2011), which investigates the contributions ofdamper flexibility to the overall response of a steel connection de-signed for damage-avoidance. Rodgers et al. (2011) develops aseries model of a damper-spring system to account for the flexibil-ity of the damper shaft and connections as well as an overallcolumn-shear versus drift connection model. Although the inves-tigation presented in Rodgers et al. (2011) is for a steel connection,the sources of flexibility associated with the dampers are very sim-ilar to those in the research presented within this manuscript.

Observed damage to the specimen was minimal. Flexural crackswere detected in the beams spaced at approximately 250 mm and amaximum width of approximately 1 mm but closed after testing.No flexural cracks were observed in the column. Up to 2% drift,virtually no cracking was observed in the joint region. Some small

cracks, approximately 50-mm in length and less than 1-mm inwidth, were observed in the beam corners from the end of the chan-nel’s flange. They formed when this region was in compressionfrom connection opening. Beyond 2% drift, some additional crackswere observed but were minor. No cracks were observed around thearmoring in the column, nor were any diagonal shear cracks ob-served across the joint. No crushing was observed around the steelarmoring the column. Upon the completion of testing, a prestressloss of 4% per thread-bar in the east-west and north-south directionoccurred.

Results from bidirectional testing are given as individual east-west and north-south direction plots in Fig. 6, with the bidirectionalclover-leaf displacement-controlled loading pattern shown inset inFig. 6(a). The specimen exhibited stable and highly dissipative hys-teresis loops over multiple cycles of different sizes with negligiblestiffness or strength degradation. The bidirectional rocking causedsome additional 50–100 mm minor crack propagation to the beamsnear the joint interface. These cracks appeared when the specimenwas displaced concurrently in the east-west and north-south (diago-nal) direction, resulting from significant force concentration at therocking corner of the beams. Comparing the uni- and bidirectionalhysteresis loops indicates that the specimen performed essentiallythe same as in uni-directional loading. Very minor strength lossfrom the bidirectional testing is noticeable at the 3-percent driftamplitude.

Fig. 8 shows the compression force in two of the dampers overtime, recorded after quasi-static testing to 3% drift. In the first 8 h,

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the initial 200 kN compression force dropped to ∼100 kN. After40 h, it reduced to ∼85 kN following the logarithmic decayin Fig. 8.

Quasi-Earthquake Displacement Test Results

Fig. 9 shows the experimental results for the three IDA selectedearthquakes used as QED test inputs. The force-displacement re-sponse, the bidirectional column orbit, and the displacement versustime are shown. In all cases, the specimen exhibited a good flag-shaped hysteretic energy dissipation response. Bidirectional rock-ing coupled with multiple nonuniform displacement cycles resultedin some minor stiffness degradation. This loss is evident when scru-tinizing the north-south force-displacement response, particularlyfor the 50% MCE case, which exhibited significant biaxial motioninteraction.

Damage to the specimen was minimal with only slight crackingnear the joint region observed; these cracks generally closed at theend of testing. Some prestress losses of 0–15 kN were detected.These losses were attributed to a combination of some local crack-ing/crushing of the concrete around the thread-bar anchorages andto normal tendon-duct frictional effects.

Multi-Level Seismic Performance Assessment

Ground motions representing different levels of seismic demandwere applied to a subassembly representing an exterior columnof a multistorey building. By using these test results, it is possibleto extrapolate observed damage to the whole structure and deter-mine if identified performance objectives were met. The first ob-jective deals with serviceability. Given a design level earthquake,there must be high confidence that the structure will not sustaindamage that disrupts normal function. The seismic demand for thisobjective is the 90th percentile DBE (10% probability in 50 years/475-year return period). The displacement profile had a single,large 1.6% interstory drift cycle followed by a slow reduction indisplacement. Residual drift was negligible. Observed damagefrom this level of shaking was minimal. Flexural cracks were ob-served in the beams and small (50 mm) cracks that closed aftertesting were observed in the beam’s joint region. The HF2V damp-ers performed well with some hysteretic energy dissipation on thefirst large pulse followed by near-elastic behavior. The specimendid not suffer any stiffness or strength degradation, as shown inthe experimental results of Fig. 6 and in more detail in Rodgers(2009). Given these results, the specimen satisfied the first MSPArequirement that the building should remain operational followinga DBE.

The second performance objective relates to repairability wherethere must be moderate confidence that the structure is repairablefollowing a rare earthquake. This criteria relates to the 50th percen-tile MCE (2% probability in 50 years=2;450 year return period).The displacement demand was most severe in the north-south(gravity) direction corresponding to a maximum interstory driftof 1.6%. Again, the specimen performed very well. Only a fewadditional cracks near the beam’s armoring were observed. Inthe east-west direction, the specimen behaved elastically with nostiffness or strength degradation. In the north-south direction, someminor stiffness degradation was observed and attributed to bindingof the draped thread-bar profile within the duct. Hence, the speci-men easily met the second MSPA requirement.

The final performance objective relates to collapse preventionand life-safety. With a high level of confidence the structure mustnot collapse following a very rare earthquake. This objective wasrelated to the 90th percentile MCE. The displacement demand

consisted of one primary pulse to 2.8% east-west interstory drift(3.1% radial drift) followed by several small cycles. As with theprevious two earthquakes, damage to the specimen was minimal.Some previously developed cracks propagated away from the jointanother 100 mm. A few additional cracks formed near the armoringregion but mostly closed after testing. During the first pulse, a con-siderable amount of hysteretic energy dissipation was observed,and subsequent displacement cycles were elastic with the full stiff-ness and strength of the specimen preserved. Some prestress lossesof ∼0–5% were recorded, likely attributed to embedding in theanchorage regions. These losses were deemed too small to neces-sitate restressing the thread-bars. Given the damage outcome fromthis level of demand, the structure satisfied the final objective oflife-safety.

It is considered that this dual experimental-computationalMSPA has verified the seismic resistance capability of the proposedstructural system with jointed precast concrete with internal HF2Vdampers. The investigation has demonstrated that the specimen,and indirectly the structure, is capable of remaining essentiallydamage-free under maximum considered severe ground-shakingconditions. All performance objectives related to serviceabilityand life-safety were achieved. This structure was designed to avoiddamage by rocking at specially detailed joints and, thus, offers anattractive alternative to conventional monolithic design and con-struction. The authors believe an equivalent monolithic structurewould have likely undergone severe cyclic rotations at its plastichinges resulting in major damage locally and excessive residualdisplacement of the global system.

Discussion

Overall, the specimen performed as expected by exhibiting verystable hysteresis loops with no stiffness or strength degradationand negligible residual drift. Minor cracks were observed in thejoint region of the beam ranging in length from 25 to 150 mm.These cracks were confined primarily to the armored end regionsand tended to close after testing. No damage (cracking or crushing)was observed in the column. This result was attributed to the aver-age axial stress (∼0:1f 0c) in the column that suppressed cracking.

Testing confirmed that the detailing strategy is sufficient to pro-tect the members from damage resulting from bidirectional drifts upto 4%. The joint was shown to ‘roll’ slightly, but this effect wasminimal. Given the good agreement between predicted and exper-imental results by using the rigid-body assumption and the negli-gible rolling of the connection, the hand analysis method isdemonstrated to be of sufficient sophistication for predicting thebackbone response.

The design of the beam-column joint region was relatively sim-ple. However, it is important to emphasize that checks be made toensure that armoring is sufficient for spreading the compressivecontact forces to the concrete without damage. Because very littlecracking was observed in the joint region, the steel fiber reinforcingin the closure pour did not play a critical role in the joint’s response.It is, thus, suggested that the addition of fibers to the mix are notstrictly necessary. To further investigate these types of effects, afinite-element analysis may be necessary to identify any portionsthat may have been overdesigned.

A particular focus of this research was an examination of theefficacy of the internal HF2V dampers. Results show that theHF2V dampers successfully achieved their design objectives.The HF2V dampers provided stable, repeatable energy dissipationand were shown to ‘reset’ after testing, as the residual force in thedevices decayed logarithmically. These results are very desirable

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particularly considering that the devices would not have to be serv-iced or replaced following an earthquake.

Because of the very small connection opening (∼5–10 mm), theeffectiveness of the HF2V dampers is very sensitive to their stiff-ness and connection to the structural system. To limit stiffnesslosses from connecting elements, the HF2V dampers were tight-ened against the connecting rods and the threaded rods groutedin the column. However, some minor slack in the connecting ele-ments led to a reduction in stiffness of approximately 60%. Hence,further development of these devices and systems should considerspecific connection details.

A considerable amount of effort was devoted to identify themost constructible solution possible. The key to this design isthe inclusion of the cast in situ closure pour. This approach providesa means of ensuring a reliable contact surface at the joint and easiercoupling of various elements, thus, enabling segmental precastconstruction.

The MSPA method relies heavily on the computational modeldeveloped from the experimental data. For realistic displacementprofiles to be extracted from the model, the response of the twomust be reasonably identical. Furthermore, nonstructural damage,which constitutes a large portion of overall damage, should be con-sidered for a more complete conclusion to be drawn. Nevertheless,the MSPA method has demonstrated a sound means of experimen-tally verifying the performance of structures at various levels ofseismic demand.

Conclusions

The following conclusions can be drawn from this study:1. The jointed precast concrete specimen with internal HF2V

dampers in the beam-end regions satisfied all seismic perfor-mance objectives related to serviceability and life-safety. Thespecimen remained essentially damage-free, fulfilling theDamage-Avoidance Design objective, after being subjectedto displacement profiles representing a design basis earthquakeand more severe maximum considered earthquakes

2. The HF2V dampers mounted internally at each beam-end pro-vide reliable, repeatable energy dissipation on every test cycle.Residual device compression forces were shown to logarithmi-cally decay through creep toward zero force over time. Thedevices do not need to be serviced or replaced following anearthquake.

3. The detailing strategy provided an excellent level of protectionfrom damage, although still retaining a degree of simplicityand constructability. Throughout the quasi-static and QED test-ing up to 3.1% radial drift, the specimen suffered negligibledamage, and only a few small cracks were observed nearthe steel armoring and these generally closed after testing

References

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