Seismic Behavior of Precast and Post-Tensioned Exterior Connections with Ductile Headed Rods

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ACI STRUCTURAL JOURNAL

TECHNICAL PAPER

Title No. 118-S07

Seismic Behavior of Precast and Post-Tensioned Exterior Connections with Ductile Headed Rods by Sanghee Kim, Thomas H.-K. Kang, Donghyuk Jung, and James M. LaFave

This study experimentally investigates the seismic structural performance of ductile bolted connectors and unbonded post-tensioning tendons for exterior precast concrete (PC) beam-column connections. Three full-scale beam-column subassemblies—a monolithic reinforced concrete (RC) connection conforming with ACI 318-19, a PC connection with ductile rods, and a PC connection with ductile rods and post-tensioning—were fabricated for quasi-static cyclic loading tests. The PC specimens exhibited stable lateral load resistance up to ±4% inter-story drift, with concentrated inelastic deformation within the ductile rods per the intended design concept. After reaching their peak lateral loads, the PC specimens experienced strength degradation comparable to the RC specimen, demonstrating that they are adequate for a lateral load-resisting system. The prestressing force induced by the post-tensioning was found to be effective at improving flexural and shear strength of the beam-column connection and at reducing the pinching effect observed at the beam-column interface of the PC specimens. Keywords: beam-column connection; embedded ductile rod connection; post-tensioned concrete; precast concrete; seismic performance evaluation.

INTRODUCTION Precast concrete (PC) or prefabricated concrete is often preferred for its fast construction speed (Probst et al. 2010; Ibrahim Ary and Kang 2012; Kang and Ibrahim Ary 2012; Hajdukiewicz et al. 2019; Park et al. 2019). Various innovative beam-column connection assemblies have been explored to take advantage of the quality and time efficiency of PC members (Martin and Korkosz 1982). For example, Pillai and Kirk (1981) tested nine full-scale PC beamcolumn connections welded via embedded steel plates in comparison with two monolithic reinforced concrete (RC) reference specimens. Bhatt and Kirk (1985) further investigated welded joints by improving a main weak link found in the previous research and thereby eliminating premature weld failure. In 1987, the National Institute of Standards and Technology (NIST) initiated a series of 10 tests on one-third-scale PC beam-column connections subjected to cyclic inelastic displacements. In a total of four phases, the research focused on hybrid connections containing both mild and post-tensioning (PT) steel contributing to the moment resistance of a beam-column connection (Cheok and Lew 1990, 1991, 1993a,b; Cheok and Stone 1994). The PT steel clamped the beam against the column to provide shear resistance and self-centering effects, while the mild steel provided ductility during cyclic loading. In a series of workshops by the Precast/Prestressed Concrete Institute (PCI) in 1991, a group of PC producers, design engineers, and contractors proposed several design ACI Structural Journal/January 2021

concepts for the Precast Seismic Structural Systems (PRESSS) project (Nakaki and Englekirk 1991). The spaced-out thread bar frame concept, with a bolted connection of the beam and column formed by wrench-tight threaded rods, had an advantage of the versatile arrangement of its frame with the same basic configuration. Merging this concept with hybrid connections, Nakaki et al. (1994) tested a ductile PC frame (DPCF) system using a connector to assemble the longitudinal reinforcement of the PC beam with a ductile rod embedded in the column. DPCF ensures the beam will behave within its elastic range, when connected to the column, while the embedded ductile rod is the only component that undergoes plastic deformation. Thus, unlike conventional strong column-weak beam RC moment resisting frame concepts, DPCF relocates the plastic hinge to within the joint region (Englekirk 2003). Although pinching was observed, the connection subassembly did not exhibit strength degradation until it reached more than 4% drift, and the ability of the column to support load was not compromised. As part of the PRESSS program, the University of Minnesota and the University of Texas at Austin also conducted experiments and nonlinear analyses to evaluate the behavior of various ductile connection systems. There were eight beam-column connection specimens with three behavior categories: tension-compression yielding, energy-dissipating, and nonlinear-elastic (Palmieri et al. 1996). Except for specimens (UT-GAP and UT-DB) with only an indirect path for force transfer, the other six specimens performed well. Three two-thirds-scale hybrid frame beam-column connections (one interior, one exterior, and one corner) were tested as a proof-of-concept for PC unbonded PT hybrid frames at the University of Washington (Kim 2000; Kim et al. 2004). A substantial proportion of total drift was due to the beams pulling away from the columns, but the self-centering effect of the PT tendons allowed significant drift ratios to be achieved with little damage. More recently, the experimental behavior of two full-scale interior beam-column connection subassemblies (one with a single concentric PT tendon), using ductile rods embedded in high-strength concrete columns, was summarized following the DPCF design methodology (Chang et al. 2013). The top connector ACI Structural Journal, V. 118, No. 1, January 2021. MS No. S-2019-412.R2, doi: 10.14359/51728179, received November 15, 2019, and reviewed under Institute publication policies. Copyright © 2021, American Concrete Institute. All rights reserved, including the making of copies unless permission is obtained from the copyright proprietors. Pertinent discussion including author’s closure, if any, will be published ten months from this journal’s date if the discussion is received within four months of the paper’s print publication.

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assembly was covered with cast-in-place concrete emulating the slab, while the lower connector was exposed, forming a T-beam cross-section. Most of the damage was found at the beam-column interface, as the embedded rod yielded in the joint region. Despite the numerous research and application efforts, previous PC connection approaches (that is, welded, hybrid, and PT) have been criticized for their inadequacy in field applications associated with a heavy workload on the construction site. Conventional ductile rod-type connections showed promising structural performance with high constructability, but they are considered economically unfavorable because of the high price of the ductile link connection. Therefore, there is the need for a new connection system that can satisfy various aspects including on-site constructability, structural performance, and economic feasibility. The ductile rod system proposed in the current study aims to maintain the advantages of conventional ductile rod connection systems (Nakaki et al. 1994; Englekirk 2003), while at the same time promoting economic efficiency of the connector by adopting lower-cost ductile rods. This connection system can for instance be widely used for construction of PC pipe racks in industrial plants, substituting for steel frames that require high material cost and additional on-site work such as drilling and welding. In this paper, the seismic performance of the ductile rod system for exterior PC beamcolumn connections is experimentally assessed to specifically address its application in areas of high seismicity. RESEARCH SIGNIFICANCE Despite the advantages of ductile connection systems for PC frames, only limited experiments on embedded ductile rod connections have been conducted, and previous studies have mainly focused on the behavior of interior connections. In the current research, a new connection system with embedded ductile rods is proposed for exterior beamcolumn joints. Discussions on design, fabrication, and experimental testing of full-scale exterior connections can shed light on both the practicality and seismic performance of the proposed connection system. EXPERIMENTAL PROGRAM Three full-scale exterior beam-column connection subassemblies (one monolithic RC and two PC specimens) were constructed and tested under reversed cyclic quasi-static lateral loading to assess the seismic performance of an embedded ductile rod connection system. The concept of a ductile rod connection and its design methodology are first discussed, and details of each beam-column joint specimen are then provided. The performance of the ductile rod connection is subsequently evaluated, including the effect of PT, in comparison with a traditional monolithic connection. Design concept and methodology In contrast to typical ductile RC beam-column connections that enforce a plastic hinge at the beam end, connections using ductile rods (Nakaki et al. 1994; Englekirk 2003) are designed to develop a plastic hinge within the joint by concentrating inelasticity in the ductile rods. In this type of 88

Fig. 1—Schematic plan and concept of ductile headed rod connection system. connection, high stress and strain demands that would have been imposed on the beam are instead transferred to the joint. Concrete in the joint can be effectively protected by sufficient compressive force in the column (less than 30% of its axial capacity) and/or lateral confinement provided by adjacent column ties. The use of bolt-type connections allows for easy and fast assembly of beam and column members on a construction site that only requires limited labor. Figure 1 shows the configuration of the ductile headed rod system comprising two pairs of ductile rods, steel transfer blocks, high-tension bolts, and high-strength threaded bars. The ductile rods with circular heads (for anchorage) are embedded in the column and connected to a transfer block with high-tension bolts. The transfer block acts as a connector for the ductile rods and high-strength threaded bars, and it transfers shear and bending moment between the beam and column members. Because the ductile rods are the only yielding component, the connection system must be designed such that the rest of the components have higher strength than that of the rods (and therefore behave elastically). To control the occurrence of yielding in the desired ACI Structural Journal/January 2021

region, each ductile rod has an effective yield zone with a reduced cross-sectional area (refer to Fig. 1). Equations (1) and (2) were employed in the design of the ductile rods for this study

ϕ1AboltFnt ≥ Arodfy,rod (1)



ϕ2nthreadAthreadfy,thread ≥ nrodArodfy,rod (2)

where ϕ1 is the safety factor 0.75, following the Specification for Structural Steel Buildings (AISC 2016); Abolt is the cross sectional area of a high tension bolt (mm2); Fnt is the nominal tensile stress of a high-tension bolt under combined tension and shear, 780 MPa (113 ksi) for ASTM A490 bolts (AISC 2016); nrod is the number of ductile rods; Arod is the cross-sectional area of the ductile rod (mm2); fy,rod is the yield stress of the ductile rod (MPa); ϕ2 is the strength reduction factor 0.9; and nthread, Athread, and fy,thread are the number, cross-sectional area (mm2), and yield stress (MPa) of the high-strength threaded bar, respectively. The ductile rod system can be used as a moment connection in the construction of a PC frame, as also depicted in Fig. 1. In the weak axis (Y-axis) direction of the frame, ductile rod connections can for instance be applied consecutively to every bay, while they can be applied partially (for example, one every three bays) in the strong axis (X-axis) direction in this example. The ductile rod has a separate circular head that can be assembled with the rod. The uneven surface of the circular head ensures its secure embedment in concrete and prevents unwanted rotation of the rod while it is being connected with the transfer block. ASTM A490 bolts (refer to Fig. 1) were used for the connection of the ductile rods and transfer block, with a tightening torque of 1393 N-m (Eq. (3)), inducing 25% of the nominal tensile strength of the bolts for a snugtight condition

T = K ∙ Dbolt ∙ Nbolt (3)

where T is torque (N-m); K is the coefficient of torque 0.15 (AIK 2016); Dbolt is the diameter of a high-tension bolt (m); and Nbolt is the bolt tension generated during tightening (N) (25% of the yield strength). In the proposed connection system, the shear transfer mechanism at the beam–column interface is through friction between the shim plates. The ASTM A490 bolts play a crucial role in this shear transfer mechanism. When the connection experiences alternating shear and moment under earthquake loading, one of the transfer blocks will be on the flexural tension side where the friction resistance of the shim plates may be reduced. However, the plates are compressed due to pre-tensioning of the bolts, and so sufficient friction resistance can always be maintained on their surfaces, providing stable and slip-free shear transfer. The ductile rod embedded in concrete resists pull-out force through bearing pressure acting on the circular head. For secure anchorage, the headed rod has design bearing strength exceeding its tensile strength (Eq. (4)), according to Section 22.8.3 of ACI 318-19 (ACI Committee 318 2019) ACI Structural Journal/January 2021



ϕVn = 0.85Ahead(2fc′) ≥ λArodfy,rod (4)

where Ahead is the cross-sectional area of the head (mm2); fc′ is concrete compressive strength (MPa); Arod is the cross-sectional area of the ductile rod (mm2); λ is an overstrength factor of 1.25; and fy,rod is the yield stress of the ductile rod (MPa). The connection system is assumed to reach its moment capacity when the ductile rods start to yield (Englekirk 2003). Design moment strength of the PC beam-column connection can therefore be calculated in accordance with ACI 318-19 using Eq. (5)

Mn = nrodArodfy,rod(drod – drod′)

(5)

where Mn is nominal moment capacity; nrod, Arod, and fy,rod are the number, cross-sectional area (mm2), and yield stress (MPa) of the ductile rods, respectively; and drod and drod′ are the distances from beam’s top face to the center of the bottom ductile rod and top ductile rod, respectively. In the design process, compression in the threaded bars was also checked and found to be lower than their buckling load. The design shear strength of the joint and the shear force acting on the joint can be estimated per ACI 352R-02 (Joint ACI-ASCE Committee 352 2010). In this study, all three test specimens were designed to have the same joint shear capacity. For the purpose of exploring interactions among the individual connection components and damage mechanisms at the beam-column joint, the specimens were deliberately designed to have the ratio of joint shear demand to capacity close to unity (that is, Vj/Vn ≈ 1) and to experience joint shear failure at a large drift ratio

Vn  1.0 f c Aj (6)



Vj = 1.25nrodArodfy,rod – 1.25Mn/lc (7)

where, fc′ is concrete compressive strength (MPa); Aj is the joint effective area (mm2); and lc is the length of column (m). Details of test specimens The new connection system investigated in this study was developed for exterior beam-column connections of pipe rack frames in industrial plants. It is designed to exhibit sufficient flexural deformation capacity (close to 5% inter-story drift ratio) without significant force reduction for mid-rise PC pipe racks of four to six stories located in active seismic areas. Each full-scale beam-column connection subassembly was accordingly fabricated to have a story height and span length of 3.5 m (11.5 ft.) and 8.1 m (26.6 ft.), respectively. Different design parameters were applied to the three specimens, which are hereby referred to as: 1) the conventional monolithic RC specimen (C-RC); 2) the PC specimen with embedded ductile rods (DRPC); and 3) the PC specimen with embedded ductile rods and PT (DRPC-T). All specimens had column and beam cross-sectional dimensions of 500 x 500 mm (19.7 x 19.7 in.) and 400 x 650 mm (15.7 x 25.6 in.), respectively. The specimens were designed and constructed 89

with concrete having a nominal compressive strength of 35 MPa (5 ksi) and steel reinforcing bars with a nominal yield strength of 400 MPa (60 ksi). Details of the test specimens are given in Table 1 and Fig. 2. C-RC was a monolithic beam-column connection specimen constructed without cold joints (used as a control specimen in the study). It was designed based on a strong column-weak beam philosophy, satisfying requirements of ACI 318-19 and ACI 352R-02 for a special moment frame. Five D29 (db = 29 mm [1.14 in.]) bars were used as longitudinal reinforcement at the top and bottom of the beam, and closed rectangular D10 (db = 10 mm [0.39 in.]) stirrups were provided uniformly along the beam at a spacing of 100 mm (3.94 in.). Clear cover thicknesses of concrete at the top/bottom and the side of the beam section were 40 and 25 mm (1.57 and 0.98 in.), respectively. Twelve D25 (db = 25 mm [0.98 in.]) bars were longitudinally located (four bars per face) along the column, and D13 (db = 13 mm [0.51 in.]) outer and inner ties were placed at 125 mm (4.92 in.) spacing, providing four legs acting in either direction. The clear cover thickness of concrete in the column was uniform at 25 mm (0.98 in.). DRPC was a PC beam–column joint subassembly connected with the new ductile headed rod system. For the beam, four D22 (db = 22 mm [0.87 in.]) bars and four high-strength threaded bars (db = 40 mm [1.57 in.]) were used as longitudinal reinforcement, along with D10 (db = 10 mm [0.39 in.]) stirrups placed every 100 mm (3.94 in.). The column was provided with the identical longitudinal and transverse reinforcement used in C-RC. Two steel transfer blocks (420 x 130 x 100 mm [16.5 x 5.1 x 3.9 in.]) with eight total sockets were linked to the end of the beam, which had a reduced cross section of 400 x 370 mm (15.7 x 14.6 in.). For the connection of the high-strength threaded bars with the transfer block, the hex nuts were first screwed onto each threaded bar toward the top of the beam (refer to Fig. 1). After the threaded bars were fastened into the inner sockets of the transfer block, the hex nuts were unscrewed toward the bottom end until they came into contact with the transfer block and were then firmly secured. The ductile rods embedded in the column were joined to the outer sockets through a bolted connection. The total length of the ductile rods was 300 mm (11.8 in.), including an effective yield zone that had a length of 170 (= 132 + 38) mm (6.7 in.). Dimensions of the ductile rods were carefully designed to induce yielding of the unbonded rods in a timely manner. If the rods are longer than needed, their yielding may be delayed or may not occur at all even at high drift ratios; if the rods are too short, they may yield before sufficient drifts are reached (or pull-out may occur). Between the transfer block and the column surface, round steel shim plates were placed, with a total thickness of 25 mm (0.98 in.), to secure a rigid connection for the high-tension bolts and ductile rods. Torque on the bolted connection was calculated based on recommendations by Englekirk (2003), using 25% of yield strength, and following a design manual for standard connections using high-strength bolts by the Korean Society of Steel Construction (2009). The connection assembly was exposed, without a cast-in-place closure, to observe

90

Table 1—Specimen design details and material properties Specimens

Beam

Column Joint

Material, MPa

C-RC

DRPC

DRPC-T

Size, mm

400 x 650

400 x 650

400 x 650

sbeam, mm

100

100

100

Mn, kN-m

682.9

659.8

685.6

Mpr, kN-m

831.9

824.8

834.2

Size, mm

500 x 500

500 x 500

500 x 500

scol

125

125

125

Vu,joint/Vn,352

1.0

0.96

0.97

fc′,meas (fr,meas)

38.3 (3.34)

38.5 (5.54)

36.9 (3.18)

fy1

536.5

554.8

554.8

fy2

487.7

487.7

487.7

fy,thread

—

1001

1001

fy,rod

—

413.7

413.7

fu,tendon

—

—

1893

Fu,eff

—

—

775

Note: sbeam is the spacing of stirrup; Mn is the nominal moment capacity; Mpr is the the probable moment capacity considering the overstrength of the materials; scol is the spacing of column confinement; Vu,joint and Vn,352 are the shear force acting on joint and nominal shear force of joint according to ACI 352, respectively; fc′,meas and fr,meas are the measured compressive and flexural strengths of concrete, respectively; fy1 and fy2 are the yield strengths of reinforcing bar and stirrup in beam except for high-strength threaded bar, respectively; fy,thread is the yield strength of high-strength threaded bar; fy,rod is the yield strength of ductile headed rod; fu,tendon is the ultimate strength of tendon; and Fu,eff is the effective prestresing force of tendon on test day.

rotation at the beam-column interface and elongation of the embedded ductile rods. DRPC-T was a PC specimen identical to DRPC except that PT tendons were added to assess the possibility to increase negative moment capacity at the beam end. Three seven-wire unbonded PT strands were eccentrically placed (120 mm [4.7 in.] up from the beam center) along the beam and through the column. The diameter of each tendon was 12.7 mm (0.5 in.), and its nominal tensile strength was 1860 MPa (270 ksi). The gap at the beam-column interface of DRPC-T was grouted with high-strength mortar (80 MPa [12 ksi]) for effective transfer of compression after PT. Due to a relatively short prestressing length (compared to practical applications), as well as shrinkage and creep of concrete, the effective prestressing force reported on the day of testing was only 775 MPa (112 ksi), approximately 41% of the measured ultimate stress of the tendon. Test setup and loading plan Figure 3 schematically illustrates the cyclic loading test setup configuration for the T-shaped beam-column connection specimens. The column was placed horizontally and pin-supported at both ends, which were assumed as inflection points. An actuator was installed at the top of the (vertical) beam to apply (cyclic) lateral loading. The loading protocol, in conformance with ACI 374.1-05 (ACI Committee 374 2014), is presented in Fig. 4. Displacement-controlled loading was applied quasi-statically (0.3 to 1.5 mm/s [0.01 to 0.06 in./s]), with gradual increase in target inter-story drift ACI Structural Journal/January 2021

Fig. 2—Specimen drawing. (Note: Units in mm; 1 mm = 0.039 in.) ratios from 0.5 to 5% and repetition of three loading cycles at each target drift. It was assumed that the test specimens reached their ultimate deformation when they experienced a certain level of force reduction. Testing continued until the lateral load decreased by more than 20% of the maximum measured force. Axial compression was not applied to the column to conservatively assess connection performance; the relatively low actual compressive force in PC columns can be beneficial in securing the ductile rods and reducing concrete damage. Global and local measurements were monitored using various instruments including linear variable differential transformers (LVDTs), a string potentiometer, load cells, and strain gauges. LVDTs were used to measure lateral displacement of the beam at the top, joint shear distortion, and flexural deformation of the beam end, as well as any slip of the specimen at the column supports. The potentiometer measured lateral displacement of the beam at midspan. Strain gauges were attached to longitudinal and transverse steel reinforcement near the joint. Load cells were placed at the column supports to monitor axial force in the column; for DRPC-T, six additional load cells were used to measure the prestressing force in the three tendons. Detailed locations of the sensors are indicated in Fig. 3.

ACI Structural Journal/January 2021

Fig. 3—Test setup. (Note: Units in mm; 1 mm = 0.039 in.) EXPERIMENTAL RESULTS Observed behavior and damage mechanisms All specimens sustained the full loading history (presented in Fig. 4) up to ±5% drift and showed stable cyclic response. Because the specimens were designed with the Vj/Vn ratio close to 1, most of the damage was concentrated around the joint region, as expected, and all of the connections 91

eventually failed in shear at the joint. Beam and column members also experienced minor to moderate damage in regions adjacent to the joint. Figure 5 presents the damage status of the specimens around the joint region at drift of 2% and 5% (final loading cycle). C-RC initially exhibited flexural and shear cracks in the beam and joint, respectively. As the drift ratio increased, shear cracks predominantly developed in the joint, forming a large X-shaped diagonal crack pattern. In DRPC and DRPC-T, initial concrete cracking appeared on the beam near the beam–column interface (refer to Fig. 5). An overlay end peeling crack developed between the top and bottom transfer blocks along the path of the connecting bars (refer to Fig. 5(b)). As the specimens were loaded cyclically, one side of the crack opened while the other side closed, and then vice versa. After initial cracking near the beam-column interface, both DRPC and DRPC-T started to accumulate shear cracks in the joint panel, as in the case of C-RC. However, unlike

Fig. 4—Loading protocol.

C-RC, which showed a large and symmetric diagonal crack pattern, the cracking in DRPC and DRPC-T was concentrated in the half of the joint nearest to the beam, forming a trumpet-shaped failure surface (refer to Fig. 5). Cracking propagated diagonally from the head of each embedded rod toward the beam, with damage clustered in the area confined by the heads of the rods. Based on this observed crack pattern, it is believed that the joint damage of DRPC and DRPC-T was strongly governed by pullout behavior of the rods. The different damage/failure patterns of the three specimens clearly demonstrate the distinct resistance mechanism of each connection system. In the monolithic system, due to its greater continuity, shear and flexural demands were fully transferred through the entire beam-column interface and resisted by the whole joint region. In the PC system, however, forces were transferred mainly through the steel blocks at the top and bottom of the beam section, with flexural resistance generated by a couple of forces acting in the ductile rods. Flexural and drift capacity Figure 6 compares the global hysteretic response of the three specimens represented by beam-end moment versus inter-story drift. The moment at the beam-column interface was calculated by multiplying the measured actuator force by the distance between the face of the column and the loading point. The inter-story drift ratio was estimated by dividing the lateral displacement of the beam at the loading point by the distance from there to the column centerline. Lateral slip of the specimens, measured at both pinned supports, was found to correspond to less than 0.05% drift and thus was simply ignored. Table 2 presents primary test results obtained from the three specimens and compares them with corresponding expected values.

Fig. 5—Damage patterns observed from specimen joints. 92

ACI Structural Journal/January 2021

Fig. 6—Moment-drift relationship. (Note: 1 kN-m = 0.738 kip-ft.) Table 2—Summary of test results Specimen C-RC DRPC DRPC-T

Dir.

Mn,cal, kN-m

My, kN-m

Mpeak, kN-m

Mpeak/ Mn,cal

Vj,peak, kN

Vn,352, kN

Vj,peak/ Vn,352

Δy, %

Δpeak, %

Δ0.8, %

u

+

894.9

825

918

1.03

1480

1393

1.06

1.72

2.88

4.97

2.9

–

894.9

840

925

1.03

1491

1393

1.07

1.65

2.99

4.8

2.9

+

657.8

625

713

1.08

1223

1396

0.88

2.6

3.91

—

1.92*

–

657.8

610

678

1.03

1162

1396

0.83

2.69

4

4.24

1.57

+

768.3

770

859

1.12

1415

1367

1.04

2.77

3.98

—

1.81*

–

687.5

675

752

1.09

1239

1367

0.90

2.73

3.96

—

1.83*

At least.

*

Note: Mn,cal is the calculated peak moment based on measured material properties; Mpeak is the measured peak moment; Vn,352 is the calculated shear strength at joint based on ACI 352R-02 (2010); Vj,peak is the measured peak shear strength of joint (Vj,peak is measured peak beam moment/0.9d – Vcol, where d is the effective beam depth and Vcol is the column shear); My and Δy are the yield moment and drift at yielding point estimated based on ACI 374.2R-13 (2013), respectively; Δpeak is the drift ratio at peak moment; Δ0.8 is the drift ratio at 80% level of Mpeak during the post-peak; and u is the ductility (Δ0.8 divided by Δy).

Although the three specimens were expected to have comparable flexural strength, C-RC recorded much higher moment capacity compared to the PC specimens, as indicated in Fig. 6 and Table 2. This is attributed to the unexpectedly high tensile strength of the beam longitudinal bars used in C-RC. From material testing performed after design of the test specimens, yield strength of the D29 bars turned out to be 536 MPa (77.7 ksi), one-third higher than their nominal yield strength; measured ultimate tensile strength was approximately 698 MPa (101.2 ksi). The moment capacity (Mn,cal) calculated based on the updated yield strength was quite close to the actual peak moment (Mpeak). C-RC yielded at drift ratios of +1.72% and –1.65%, and recorded its peak moment at a drift ratio of 3% in both directions. C-RC was able to maintain a high level of flexural resistance (comparable to its peak moment) during drift ratios of 2to 4%, and 80% of the peak moment (or above) until it reached 5% drift, resulting in a displacement ductility of at least 2.9. In the case of DRPC, the measured moment capacity matched quite well with the expected value because yield strengths of the ductile rods and high-strength threaded bars were quite close to their nominal strengths. Yielding of DRPC was reported at drift ratios of +2.60% and –2.69%, approximately 1% later than in C-RC. DRPC recorded its peak moments of +713 kN-m (526 kip-ft.) and –678 kN-m (500 kip-ft.) at 4% drift in both directions; these drifts satisfied a drift ratio limit of 3.5% specified in ACI 374.1-05. ACI Structural Journal/January 2021

The hysteretic response of DRPC in Fig. 6 also reveals that it experienced pinching near the center point of the graph. This pinching appears to be attributable to slip—the relative motions that inevitably occur between components of the connector and at the concrete-connector interface. The higher drifts of DRPC at yielding and peak moments, compared to C-RC, are thought to be related to the pinching effect. At drifts of ±5%, DRPC exhibited flexural resistance corresponding to 87% and 72% of the peak moment, respectively, in the positive and negative directions. Overall, the hysteretic behavior of DRPC-T was similar to that of DRPC. DRPC-T also reached its peak moments at 4% drift and exhibited apparent pinching during reversals of loading. Due to the eccentric PT tendons, DRPC-T was expected to have unequal moment capacities by developing 12% higher moment in the positive direction (refer to Table 2). As highlighted in Fig. 6, the three PT tendons indeed contributed to enhancing flexural resistance of this beam–column connection, which achieved 14% higher peak negative moment than peak positive moment. Furthermore, compared to DRPC, DRPC-T had 20% higher peak negative moment (in the positive direction) due to the PT tendons. Based on the experimental results, it is expected that eccentric PT tendons can be effectively used to resist high-gravity-induced negative moment at the beam ends, and their effect will be further enlarged as the amount of PT is increased. It is also worth noting that DRPC-T showed 93

Table 3—Degradation of measured moments normalized by peak negative moment in positive direction Specimen Cycle Drift ratio

C-RC

DRPC

DRPC-T

First

Second

Third

First

Second

Third

First

Second

Third

4%

0.97

0.91

0.82

1.00

0.96

0.80

1.00

0.91

0.85

5%

0.80

0.63

0.54

0.87

0.73

0.67

0.85

0.71

0.62

Fig. 7—Shear distortion behavior of joint panel. higher initial resistance (slope) in the positive direction in Fig. 6, somewhat alleviating the pinching effect. This observation agrees well with previous test results (Chang et al. 2013), and reveals that the restoring force of the prestressed tendons can be effective in reducing pinching. When subjected to drifts of ±5%, DRPC-T still retained 85% and 92% of the peak moments in the positive and negative directions, respectively. To assess post-yield stability of the specimens under repeated loading, degradation of lateral resistance at the high drift ratios was investigated. In Table 3, negative moments measured at 4% and 5% drifts are normalized by the peak negative moment, and their degradations at each loading cycle are compared. At 4% drift ratio, all specimens exhibited similar levels of degradation, with the normalized moments as low as 0.80. At 5% drift ratio, the normalized moment of C-RC dropped to 0.54 at the third cycle, but DRPC and DRPC-T maintained relatively higher residual moments at the same cycle, of approximately 0.58 and 0.62, respectively. The post-yield stability of the PC specimens was similar to or greater than that of C-RC. Joint shear distortion Figure 7 compares shear distortion behaviors of the beam–column joints. Shear distortions were measured by the LVDTs at the joint panel up to 4% drift (before failure of the specimens). The joint shear force demand (Vj,test) on the vertical axis of each graph is normalized by the nominal joint shear strength (Vj,352) estimated in accordance with ACI 352R-02 (refer to Table 2). As mentioned earlier, a target normalized shear force (Vj,test/Vn,352) of 1 was applied in the design of all specimens. In C-RC, the peak shear force acting on the joint was quite close to its design strength, leading to Vj,test/Vn,352 of approximately 1.1 in both directions (refer to Table 2). The joint panel of C-RC was distorted at angles from +0.006 rad to –0.007 rad under drifts of ±3%, respectively. Joint distortion further 94

increased to as high as –0.011 rad at a drift of –4%; joint shear distortions of approximately 0.01 rad have typically been associated with reaching an RC joint’s shear strength in well-detailed normal-strength concrete exterior connections (Kim and LaFave 2012). The peak shear force in DRPC was reported as 88% and 83% of the design strength in the positive and negative directions, respectively. Although the joint panel still had an estimated safety margin of more than 10% for overall joint force, it sustained critical shear damage triggered by pullout of the ductile rods. DRPC recorded distortion angles ranging from +0.008 rad to –0.006 rad at drifts of ±3%; the distortions doubled to +0.014 rad to –0.017 rad at ±4% drift. These increased distortions were accompanied by the development and widening of trumpet-shaped shear cracks in the joint panel, which in turn weakened anchorage of the ductile rods. It is for some of these same reasons that it has been suggested for exterior connections that a lesser effective joint depth should be used in nominal joint shear strength calculations (AIJ 1999), to better reflect the actual joint mechanics. Compared with DRPC, DRPC-T showed a similar pattern of joint shear behavior in that it recorded distortion angles from +0.007 rad to –0.006 rad between ±3% drift ratios, and the angles greatly increased to +0.013 rad to –0.012 rad as the drift ratio increased to ±4%. However, DRPC-T recorded slightly higher shear force corresponding to 104% and 90% of the design strength in the positive and negative directions, respectively. Also, the overall variation range of distortion angles was smaller than that of DRPC. The higher shear force and lower shear distortion of DRPC-T are thought to be the effect of the PT tendons, which continuously provided confining force to the joint panel. Stiffness degradation Based on the cyclic response of the three specimens provided in Fig. 6, moment-drift envelope curves are drawn and compared in Fig. 8, in which the measured moment was ACI Structural Journal/January 2021

Fig. 9—Secant stiffness. Fig. 8—Comparision of envelope curves. normalized by the peak moment in each direction. Among the three specimens, C-RC (which had the monolithic connection) showed the highest overall response slope, followed by DRPC-T and DRPC. Overall, both PC specimens exhibited comparable response slopes, but DRPC-T had a slightly higher slope at the initial loading phase because of the PT. Slopes of the PC specimens can be better compared through a close-up graph displaying the center region of the entire response (Fig. 8 inset). In this graph, it is seen that the negative envelope curve of DRPC starts at a drift ratio of +0.11%, which indicates residual deformation generated from the load (0.5% drift) applied in the positive direction. In DRPC-T, the negative envelope curve starts at the origin, with no residual deformation, due to the self-centering effect of the PT. Also, the graph clearly shows that lateral deformations of the PC specimens were substantially affected by slip occurring at the jointed connection. To assess the influence of slip on lateral response of the specimens, secant stiffness at the first loading cycle was compared at different levels of drift in Fig. 9. For the PC specimens, two stiffness parameters (K1 and K2) were defined considering the effect of slip. In Fig. 8, the drift ratio that exhibited low lateral resistance due to slip was denoted as Δs. K1 indicates the stiffness starting at the origin of the graph without taking into account Δs, while K2 is the one starting at the end of Δs. In Fig. 9, K1 of DRPC and DRPC-T increased at drift ratios below 1% in both directions, in contrast to C-RC that showed continuous secant stiffness degradation, indicating that the contribution of slip to lateral displacement of the PC specimens was substantial in the early phase of loading. K2 exhibited similar trends to K1, but of a much lower magnitude. From drift of 1.5%, DRPC and DRPC-T started to show decreasing secant stiffness, which later approached that of C-RC at drifts above 3%. The experimental results in Fig. 9 demonstrate that C-RC initially showed much higher lateral stiffness than those of the PC specimens, up to drifts of ±2%, due to excellent steel-concrete bond at the beam-column interface. As the drift ratio further increased, though, the stiffness gap between C-RC and the PC specimens decreased, and they showed comparable lateral stiffness at drifts above 3%. Compared to the monolithic connection, the ductile connection system may have higher displacement demands due to the lower initial lateral stiffness. At drift ratios above 3%, however, the difference in lateral stiffness of the two systems is almost ACI Structural Journal/January 2021

Fig. 10—Reduction of lateral stiffness under repeated loading. negligible, and its impact on the displacement demand also becomes insignificant. Another important aspect to point out is the reduction in lateral stiffness under repeated loading. Figure 10 describes how much the lateral stiffness reduced at each target drift by presenting the ratio of secant stiffness at the last (third) cycle to that at the first cycle at the same drift. In the positive direction, all specimens exhibited relatively similar reduction rates, recording normalized stiffnesses as low as 0.66 at a drift of 5%. Normalized stiffnesses in the negative direction were maintained close to or above 0.9 initially, but after –3% drift they began to drop (at different rates). At a drift of –5%, DRPC recorded the lowest normalized stiffness of 0.58, while DRPC-T and C-RC still retained a normalized stiffness above 0.7. Compared with C-RC and DRPC-T, which had relatively high connection resistance at the beam-column interface due to the help of a monolithic connection and PT, respectively, DRPC relied solely on anchorage of the ductile rods and thus appeared to experience more severe stiffness reduction under repeated loading. Energy dissipation Hysteretic energy dissipated under cyclic loading is presented in Fig. 11. Dissipated energy was estimated by calculating the area enclosed by the hysteresis loops of the three cycles at each target drift. According to Fig. 11, C-RC showed superior energy dissipation capacity, especially from drifts of ±2%, dissipating more than twice as much energy as did the PC specimens. This difference in energy dissipation mainly originated from the different connection mechanisms of the two (that is, monolithic versus jointed) systems. It is considered that the full cross section of the 95

Fig. 11—Energy dissipated by test specimens. Fig. 13—Equivalent viscous damping ratios.

Fig. 12—Energy dissipation ratios of third to first cycle. beam at the beam-column interface (refer to Fig. 2) and the excellent steel-concrete bond of C-RC played a crucial role in its higher energy dissipation. DRPC and DRPC-T exhibited comparable energy dissipation capabilities up to drifts of ±4%, but DRPC-T dissipated more energy than DRPC at drifts of ±5%. The moment capacity of DRPC rapidly decreased at drift ratios of ±5%; however, the PT helped DRPC-T maintain relatively higher moments in its postpeak response and enabled it to dissipate more energy. Figure 12 presents relative energy dissipation ratios between the first and third load cycles. The energy dissipated by the third cycle was divided by that of the first cycle. In C-RC and DRPC, overall normalized energy was maintained at around 60% or slightly less on average. DRPC-T initially showed higher normalized energy (even greater than 1 in the positive direction), probably because of the PT effect. However, it decreased rapidly and became comparable to those of C-RC and DRPC from 1.5% drift ratios. For further comparison of energy dissipation, the equivalent viscous damping ratio, ξeq = (2Aloop/πArect), was also considered, where Aloop indicates the energy dissipation (enclosed area) of the first cycle, and the area of circumscribed rectangle of the hysteresis loop is denoted as Arect (Lim et al. 2018). Figure 13 shows the damping ratios of the three specimens at each target drift ratio. All three specimens exhibited similar trends at drift ratios below 2%. At the drift ratio of 0.5%, the damping ratios of C-RC, DRPC, and DRPC-T were calculated as 6.9%, 8.3%, and 4.9%, respectively. Those values appear reasonable and correspond well with typical damping ratios (5%) of RC structures in the cracked elastic range (fib 2013). Up to drift ratio of 1%, the 96

Fig. 14—Average prestressing force of tendons. damping ratios of the specimens reduced to as low as 3.2%. The damping ratios started to increase from 2% drift ratio. Among all specimens, C-RC recorded the highest damping ratio (15%) at 4% drift ratio. The maximum damping values for DRPC and DRPC-T were 8.2% and 7.8%, respectively, at 5% drift ratio. It is seen that the equivalent viscous damping ratios of ductile rod and monolithic connections were comparable in the elastic range. After yielding (Δy), however, the maximum damping ratio of the monolithic connection was twice as high as those of the ductile rod connections. PT behavior The prestressing force applied to DRPC-T was continuously monitored with load cells. Figure 14 presents an average prestressing force (from all six load cells) of the three eccentrically placed tendons (refer to Fig. 2) under cyclic loading. Before application of lateral loading, the average prestressing force was 79 kN (17.8 kip), and it increased to 132.5 kN (29.8 kip) as DRPC-T was deformed by 5% drift in the positive direction. At this maximum prestressing force, the tendons reached about 67.5% of their measured tensile strength (1893 MPa [274 ksi]) and contributed 18.1% of DRPC-T’s maximum moment capacity. The contribution of the tendons satisfied Section 18.6.3.5(c) of ACI 318-19, where the contribution of prestressing steel is limited to no more than 25% of the flexural strength at the critical section. Figure 14 also shows that the unbonded prestressing force ACI Structural Journal/January 2021

Fig. 16—Normalized strains of transverse reinforcement. demonstrate that inelastic deformation was well controlled within the designated areas according to the intended design concept.

Fig. 15—Normalized strains of longitudinal reinforcement. was always maintained above its initial value, even under loading applied in the negative direction, as the tendons are unbonded and there was permanent joint damage at 4% drift and larger. Behavior of longitudinal reinforcement in beam Local behavior of the longitudinal reinforcement was analyzed to better understand the flexural behavior of the test specimens. Figure 15 provides tensile strains of the beams’ longitudinal reinforcement measured near the critical section. The strain gauges were attached on the D29 bars for C-RC and on the threaded bars and ductile rods for DRPC and DRPC-T, as depicted in Fig. 15. Maximum tensile strains for each type of reinforcement were represented with respect to the corresponding yield strain (that is, yielding occurs at a normalized strain of 1). Bar strains of C-RC already exceeded the yield strain at drifts of ±0.75% and reached almost five times the yield strain at drifts of ±2%. On the other hand, the threaded bars of both DRPC and DRPC-T recorded maximum normalized strains of approximately 0.7 at ±4% drift and remained below the yield level throughout testing. Instead, both PC specimens experienced yielding of the ductile rods before they reached peak lateral strength. Yielding of the rods occurred at 1% drift or slightly more in the positive direction and at around 3% drift or below in the negative direction. In the design phase, the D29 longitudinal bars and the ductile rods were designated as the yielding components of the C-RC and PC specimens, respectively. The strain measurement data clearly

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Behavior of transverse reinforcement in joint Deformations of transverse reinforcement in the columns near the joint region can be seen in Fig. 16, where the A-, B-, and C-layers are located at different section heights (refer also to Fig. 15). A-layer and C-layer indicate the inner and outer steel ties placed parallel to the beam, respectively, and B-layer refers to those placed perpendicular to the beam. Maximum tensile strains at each target drift were normalized in relation to the corresponding yield strain. The inner ties on A-layer of C-RC yielded soon after drifts of ±1%, much earlier than those of DRPC and DRPC-T. Previous research has shown that when transverse joint reinforcement reaches its yield strain, it is a sign that joint shear capacity is limiting maximum response of overall subassembly behavior (Kim and LaFave 2007). The A-layer ties of C-RC appeared to be strongly affected by a wide tension/compression field spread out over the whole joint panel, which later developed the large X-shaped diagonal crack (refer to Fig. 5). The PC specimens showed relatively lower inner tie strains compared to C-RC. It is suspected that the A-layer ties of the PC specimens might not have been fully engaged in tension because of the concentrated joint damage in the column region nearest to the beam. Only DRPC-T experienced initial tie yielding, just after the drift ratio of +3%. It is possible that steel tendons penetrating the joint induced diagonal tension in the panel zone just like the longitudinal bars of C-RC did. Therefore, higher force demand would have occurred in the C-layer ties of DRPC-T as the beam was loaded in the positive direction. In the case of the B-layer ties, yielding was reported at +3% drift from DRPC and at –3% drift from DRPC-T. C-RC recorded a maximum normalized B-layer tie strain close to 1 at +4% drift. The higher B-layer tie strains of the PC specimens are considered to be associated with the behavior of the ductile rods, which can impose a large force demand on the ties while being pulled out.

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Fig. 17—Applications of ductile headed rod system to pipe rack structures. Field applications Development and testing of the ductile rod exterior connection system were undertaken for its application to pipe rack frames in industrial plants. In building construction, more than two beams are orthogonally connected to one column, causing congestion or interference of ductile rods in the joint (both exterior and interior). In industrial structures, however, the situation can be avoided by connecting the beams at slightly different heights or using ductile rod exterior connections only (refer to Fig. 1). Currently, this kind of exterior connection system has been used in pipe rack constructions for the: 1) Residue Fluid Catalytic Cracking (RFCC) unit project in Cilacap, Indonesia; 2) Petrochemical Plants project in Rayong, Thailand; and 3) Uzbekistan Gas to Liquid (UZGTL) project in Kashkadarya, Uzbekistan. Figure 17 presents pictures of the PC pipe racks constructed in those projects. For Cilacap RFCC project, a total of 1040 ductile rod exterior connections were used to assemble 190 columns and 1662 beams. Compared to steeltype construction, the PC pipe rack was completed 4 months faster, which cut down construction costs by 20%. Rayong Petrochemical Plants project used 434 ductile rod exterior connections for 94 columns and 683 beams, achieving a 98

2-month construction period shortening and 25% cost reduction, compared to steel-type construction. As demonstrated in these recent projects, the ductile rod exterior connection system enables faster on-site construction through easy assembly of PC members and brings important economic benefits in comparison to conventional steel pipe racks. SUMMARY AND CONCLUSIONS In this study, a series of cyclic loading tests was conducted on three full-scale exterior beam-column connection subassemblies to explore the seismic performance of a new ductile headed rod connection system. Two identical PC specimens were fabricated using the ductile rods; one of them was provided with PT tendons. For comparative study, a monolithic RC specimen was also included in the testing. Important findings from the current experimental study can be summarized as follows: 1. To achieve ductile flexural response of the PC beamcolumn connection, ductile rods placed in the joint were designed to yield and accommodate high inelastic deformations. Test results showed that yielding was concentrated in the ductile rods for both DRPC and DRPC-T, and high-strength threaded bars remained elastic. Furthermore, ACI Structural Journal/January 2021

expected flexural strength of the PC connections was accurately achieved, proving that specimen design and fabrication procedures were successfully carried out as intended. 2. The peak moment of C-RC was reported at 3% interstory drift, whereas the PC specimens recorded the peak moment at 4% drift. 3. The slip behavior was considered as an unavoidable characteristic of the PC specimens assembled with a bolted connection, but it did not adversely affect the maximum seismic capacity. Furthermore, the amount of slip occurring in the current tests is considered acceptable in non-serviceability-sensitive structures such as pipe racks in plants, warehouses, or parking structures. Cyclic response of DRPC-T demonstrated that some of this slip behavior and resulting pinching can be alleviated by introducing a PT force. 4. After exhibiting peak moment, the PC specimens then experienced significant joint damage with trumpet-shaped shear cracks. This crack pattern was closely related to the pullout behavior of the ductile rods, which eventually led to considerable strength degradation at 5% drift. If the joint panel is provided with sufficient shear capacity, such PC specimens are expected to show much more stable post-peak response. Additional research is needed to better investigate the seismic performance and damage pattern of PC beamcolumn connections with high shear capacity. 5. Advantages of the ductile rod connection system have been demonstrated through recent field applications. It can greatly reduce the construction period, with its high constructability, and save approximately 20 to 25% of the construction cost that would have been required for conventional steel pipe racks. AUTHOR BIOS

ACI member Sanghee Kim is an Assistant Professor at Kyonggi University, Suwon, Korea. He received his BS from Jeonbuk National University, Jeonju, Korea, and his MS and PhD from Seoul National University, Seoul, Korea. His research interests include the structural behavior and impact resistance of reinforced and steel fiber-reinforced concrete structures. Thomas H.-K. Kang, FACI, is a Professor of structural engineering, Director of Engineering Education Innovation Center at Seoul National University, and Adjunct Professor teaching the summer post-tensioned concrete structures course at the University of Illinois at Urbana-Champaign, Urbana, IL. He is a member of ACI Subcommittee 318-T, Structural Concrete Building Code—Post-Tensioned Concrete; Joint ACI-PTI Committee 320, Post-Tensioned Concrete Building Code; Joint ACI-ASCE Committees 352, Joints and Connections in Monolithic Concrete Structures, and 423, Prestressed Concrete; and Joint ACI-ASME Committee 359, Concrete Containments for Nuclear Reactors. He received the ACI Wason Medal for Most Meritorious Paper in 2009. His research interests include the design and behavior of reinforced, prestressed, and post-tensioned concrete structures. ACI member Donghyuk Jung is an Assistant Professor at Pusan National University, Busan, Korea. He received his BS from Korea University, Seoul, Korea; and his MS and PhD from the University of Illinois at Urbana-Champaign. His research interests include the design of reinforced concrete structures using advanced materials such as shape memory alloys. James M. LaFave, FACI, is Professor, Associate Head, and Director of Undergraduate Studies and Director of CEE Education Programs at Zhejiang University, Department of Civil and Environmental Engineering (CEE) at the University of Illinois at Urbana-Champaign. He is a member and past Chair of Joint ACI-ASCE Committee 352, Joints and Connections in Monolithic Concrete Structures, and he received the ACI Wason Medal for Most Meritorious Paper in 2009. His research interests include the behavior and design of structural concrete beam-column connections.

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ACKNOWLEDGMENTS

The work was funded by National Research Foundation of Korea (No. 2015R1A5A1037548) and Engineering Research Institute of Seoul National University. Pre-tension bolt rigid connection (PBRC) specimens were provided by Euro Engineering Co., Ltd. in Korea. The authors would like to acknowledge the engineers of Euro Engineering Co., Ltd. for their helpful discussion and former graduate students D. J. Lee at Seoul National University for her hard work on structural testing and J. H. Shin at the University of Illinois at Urbana-Champaign for his participation in data reduction. The views expressed are those of the authors and do not necessarily represent those of the sponsors or other participants.

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