Seismic performance of emulative precast concrete beam–column connections with alternative reinforcing details

Abstract

This article presents an experimental investigation on alternative reinforcing details for the bottom bars of precast concrete beams at cast-in-place beam–column joints to achieve the behaviour as for monolithic reinforced concrete beam–column connections. To relieve steel congestion and fabrication difficulties, it is proposed to use headed bars for the bottom bars that are protruded from precast beams and anchored in the middle of the beam–column joint. In total, six interior beam–column connection specimens were tested under reversed cyclic loading. The primary test variables were the transverse beams and the anchorage of the bottom beam bars in the joint. Hysteretic behaviour, including strength degradation, stiffness degradation and energy dissipation, was evaluated in accordance with the acceptance criteria for special moment-resisting frames. Test results demonstrated that emulative precast concrete specimens with bottom beam bars anchored in the joint middle can perform as well as monolithic beam–column connections with continuous beam bars passing through the joint. On the basis of the experimental results, design recommendations are drawn for these types of emulative precast beam–column connections.

Keywords

anchorage beam–column joint headed reinforcement precast concrete seismic testing

Introduction

Because of their advantages in time and cost savings, precast concrete building structures are becoming more and more common around the world. For earthquake-resistant concrete frames, several alternative solutions have been developed for moment-resisting connections between precast beam and column components to achieve the behaviour as for monolithic reinforced concrete structures (Chen et al., 2012; Choi et al., 2013; Guan et al., 2016; Ha et al., 2014; Im et al., 2013; Pampanin, 2005; Park, 2002; Watanabe, 2007; Xue and Zhang, 2014). Figure 1(a) illustrates one of the traditional arrangements for emulative precast concrete moment frames that is used in New Zealand, Japan and Taiwan. The semi-precast beam elements are placed on the edges of the column top, followed by reinforcement fabrication and concrete casting in the cast-in-place joint, beam tops and slabs. Afterwards, the upper precast column element is connected by longitudinal column bars that protrude into mortar-grouted coupling sleeves within the column to form an emulating monolithic reinforced concrete frame. This typical arrangement has many advantages such as reductions in site formwork and labour, increased construction speed and high-quality control. The primary drawback of this arrangement is the steel congestion in the cast-in-place joint, where the protruded bottom bars from each precast beam element need to be staggered well in the joint core to prevent them from conflicting with each other, as shown in Figure 2.

Figure 1.

Precast concrete beam–column connections emulating monolithic reinforced concrete frames: (a) conventional hooked bars and (b) alternative headed bars.

Figure 2.

Staggered hooked bars in an interior beam–column connection.

This article proposes alternative reinforcing details for the bottom beam bars to relieve the steel congestion in the cast-in-place connections for special moment frames. As shown in Figure 1(b), the bottom beam bars of the precast beam elements are terminated by anchor plates or heads with anchorage lengths not exceeding one-half of the column depth. Heads allow the bars to be developed in a shorter length than required for standard hooks (Thompson et al., 2006). Heads also liberate the space occupied by the bent-up hook extensions in the joint. The use of headed bars that terminate in the middle of the joint, rather than hooked bars that extended to the far side of the joint core, can relieve the steel congestion and construction difficulties in the cast-in-place connection. However, headed bars with a shorter anchorage length may also increase the potential for concrete breakout failure, which should be precluded by providing adequate reinforcement in the joint.

To avoid steel congestion of the bottom beam bars, the termination of the bottom bars of precast beams with bent-up hooks in the middle of the joint was experimentally studied by Japanese researchers in 1990s (Castro et al., 1991; Matsudo et al., 1992; Satoh et al., 1996). The experimental results showed that emulative cast-in-place connection performance could be achieved for anchoring bottom beam bars in the joint middle with additional confinement reinforcement in the joint. These reinforcing details can avoid steel congestion of the bottom beam bars from two opposite faces of the joint, but it can also result in another congestion problem related to joint transverse reinforcement in the joint core. To date, it is well accepted that the use of headed bars in place of hooked bars is a viable solution to steel congestion in beam–column connections and presents no significant design problems (Kang et al., 2009; Wallace et al., 1998). Chen et al. (2012) tested four beam–column connections using lap splices of headed bars protruding from beam bottom ends that extended to the far side of the confined core, while the top beam reinforcement was continuously extended through the joint. The researchers observed satisfactory and comparable performance between the emulative and monolithic concrete specimens subjected to reversed cyclic loading up to an inter-storey drift ratio of 3% or 5%.

More recently, Chiu et al. (2016) tested seven interior and five exterior monolithic reinforced concrete beam–column connections to investigate the seismic anchorage behaviour of headed bars with varying anchorage lengths and bar spacing. In the tested interior beam–column connections, the top beam bars were continuously extended through the joint, while the bottom beam bars were lap-splice or non-overlap anchored by heads in the joint. For a joint with low shear demand, the test results showed that the seismic anchorage performance of the bottom beam bars (either lap-spliced or non-overlapped in the joint) is as good as that of the top beam bars passing through the joint. However, for a joint with a design shear stress about $1.25 \sqrt{f_{c}^{'}} MPa$ , bottom beam bars with non-overlap anchoring have inferior performance than those that are lap-splice anchored in the joint. Therefore, Chiu et al. (2016) preferred to use lap-splice anchorage for bottom beam bars in joints with high design shear stresses. Notably, the specimens tested by Chiu et al. (2016) had an equal area of top and bottom reinforcements. In common design practice, the bottom reinforcement area $A_{s, bot}$ could be less than the top reinforcement area $A_{s, top}$ at the joint face for a special moment frame beam. For such case, a less critical and relatively lower flexural tension from the bottom beam bars would be anchored in the joint.

This article proposes that the bottom beam bars be non-overlap anchored by heads in the middle of the joint, as shown in Figure 1(b). This allows for more freedom of design, proportion, erection and fabrication in precast concrete elements. According to ACI 318-14 Section 18.9.2.3, precast concrete frames with alternative details that do not satisfy the prescriptive requirements shall be proven by experiments satisfying the acceptance criteria for moment frames defined by ACI 374.1-05 (ACI Committee 374, 2005). Therefore, an experimental programme is conducted to verify the proposed anchorage details for the bottom beam bars of precast beams. Different from prior research works, the anchorage performance of headed bottom bars was evaluated with a varied bottom-to-top reinforcement area ratio and a high shear stress acting on the joint.

Experimental investigation

Six large-scale reinforced concrete interior beam–column connection specimens, as shown in Figure 3, were constructed and tested at YunTech in Taiwan to investigate the use of alternative reinforcing details for ductile connections of precast beams and columns. The experimental programme was designed using a concrete compressive strength $f_{c}^{'}$ of 55 MPa and yield strengths of 420 and 490 MPa, respectively, for longitudinal and transverse reinforcements. The primary test variables were the ratio of the bottom-to-top beam reinforcement area $(A_{s, bot} / A_{s, top})$ , the anchorage of the bottom beam bars in the joint and the presence of transverse beams. The specimens were designed as part of a special moment frame in high seismic zones and followed the requirements of ACI 550.1R-09 (ACI Committee 550, 2009) and ACI 318-14 (ACI Committee 318, 2014), except for the joint shear stress and the anchorage details of the bottom bars that were investigated in this test programme.

Figure 3.

Test matrix and reinforcing details of beam–column joint specimens.

Connection design

Figure 3 and Table 1 show the element dimensions, reinforcing details, material properties, and test parameters for the test specimens. Three specimens, designated as Group A, used eight and four D25 reinforcing bars for the longitudinal reinforcement in the beam top and bottom, respectively, resulting in an unequal reinforcement ratio in the beam section. The other three specimens, designated as Group B, used six D25 reinforcing bars symmetrically for the top or bottom beam reinforcement, resulting in a reinforcement ratio of $A_{s, bot} / A_{s, top} = 1.0$ . The total amount of beam longitudinal reinforcement is 12 D25 reinforcing bars, and thus, the design shear force of the joint $(V_{u})$ , which is estimated by the method recommended by ACI 352R-02 (ACI-ASCE Committee 352, 2002), is almost the same for each test specimen

V_{u} = 1.25 f_{y} (A_{s, top} + A_{s, bot}) - V_{col}

(1)

V_{col} = \frac{(M_{pr}^{+} + M_{pr}^{-})}{(L_{b} - h_{c})} \frac{L_{b}}{L_{c}}

(2)

where $A_{s, bot}$ is the area of bottom beam bars; $A_{s, top}$ is the area of top beam bars; $f_{y}$ is the specified yield strength of reinforcement; $V_{col}$ is the column shear force in equilibrium with the probable beam moments $M_{pr}$ at the joint faces, which were determined using a bar stress of $1.25 f_{y}$ , for positive and negative bending moments, respectively; $L_{b}$ is the unit beam length (4.5 m for tested specimens); $L_{c}$ is the unit column length (3 m for tested specimens) or the equivalent storey height; and $h_{c}$ is the column depth of 500 mm. For the test specimens listed in Table 1, the value of $V_{u}$ is computed to be 2720 kN using equation (1).

Table 1.

Material properties and connection parameters.

Parameter		Specimen
Parameter		A1	A2	A3	B1	B2	B3
Concrete	$f_{c}^{'}$ (MPa)	56.8	61.9	49.2	55.6	51.2	53.2
Longitudinal beam	$b \times h$	400 mm × 500 mm			400 mm × 500 mm
	$A_{s, top}$	8-D25 (SD420^a)			6-D25 (SD420^a)
	$A_{s, bot}$	4-D25 (SD420^a)			6-D25 (SD420^a)
	$M_{pr}$	+484 kN m, −787 kN m			+631 kN m, −631 kN m
	$A_{v}$	4-D13 at 100 mm (SD490^b)
Transverse beam	$b \times h$	NA	NA	400 mm × 500 mm	NA	NA	400 mm × 500 mm
Column	$b \times h$	500 mm × 500 mm
	$A_{st}$	12-D25 (SD420^a)
	$A_{sh}$	4-D13 at 100 mm (SD490^b)
Joint	s (mm)	100	100	150	100	100	150
	$A_{sh} / s b_{c}$	0.0112	0.0112	0.0075	0.0112	0.0112	0.0075
	$A_{sh, ratio}$ ^c	1.10	1.01	1.69	1.12	1.22	1.56

$f_{c}^{'}$ : compressive strength of concrete cylinders at testing date; b: section width; h: section overall depth; $A_{s, top}$ : area of top beam reinforcement; $A_{s, bot}$ : area of bottom beam reinforcement; $A_{st}$ : total area of longitudinal reinforcement; $A_{sh}$ : total cross-sectional area of transverse reinforcement; $A_{v}$ : area of shear reinforcement; s: spacing of transverse reinforcement; $b_{c}$ : column core dimension measured to the outside edges of the transverse reinforcement composing area $A_{sh}$ ; $M_{pr}$ : probable flexural strength of the beam calculated using a bar stress of $1.25 f_{y}$ for positive and negative bending moments; NA: not available.

SD420 D25 reinforcement; bar area = 507 mm²; measured yield strength f_y = 470 MPa.

SD490 D13 reinforcement; bar area = 127 mm²; measured yield strength f_y = 498 MPa.

The provided amount of transverse reinforcement divided by the amount of transverse reinforcement required by ACI 318-14; the required amount for specimens A3 and B3 is one-half of that for other specimens because of confinement from transverse beams.

All test specimens had a square column section of 500 mm × 500 mm detailed with 12 D25 Grade 420 (specified f_y = 420 MPa) longitudinal reinforcing bars and D13 Grade 490 (specified f_y = 490 MPa) transverse hoops and ties at a spacing of 100 mm. The Grade 490 reinforcement was used to reduce the amount of transverse reinforcement for confinement of 55-MPa concrete column. According to the ACI 318 and ACI 352R-02 requirements for joints in special moment frames, the primary connection design parameters are computed as follows.

Column-to-beam flexural strength ratio

To produce yielding in the beams rather than in the columns, the columns shall be stronger than the beams framing into a joint of a special moment frame and satisfy

M_{r} = \frac{\sum M_{nc}}{\sum M_{nb}} ⩾ 1.2

(3)

where $\sum M_{nb}$ and $\sum M_{nc}$ are the sum of the nominal flexural strengths of the beams and columns, respectively, calculated at the joint faces. The nominal flexural strength of the column is 670 kN m under an axial compression of $0.05 A_{g} f_{c}^{'}$ for the test specimens. According to ACI 374.1-05, it is conservative not to apply axial load, which will be generally less than the balanced load and increase the flexural strength in the columns. However, zero axial load is unrealistic and may be too conservative for bond performance of beam bars in the joint. Therefore, a least column axial of $0.05 A_{g} f_{c}^{'}$ was conservatively used in this programme. The $\sum M_{nb}$ are 1047 and 1041 kN m for beams in Group A and B specimens, respectively. Thus, the moment strength ratio $M_{r}$ is 1.28 or 1.29 for the test connections in Group A or Group B, respectively. Yielding of the flexural reinforcement should occur in the beams first.

Joint shear demand-to-capacity ratio

To avoid premature shear failure of joints, the demand of joint shear force ( $V_{u}$ , equation (1)) shall not exceed the nominal shear capacity of the joint $(V_{n})$ times the strength reduction factor $ϕ$ of 0.85

V_{u} \leq ϕ V_{n} = ϕ γ \sqrt{f_{c}^{'}} A_{j}

(4)

where $γ \sqrt{f_{c}^{'}}$ is the nominal shear stress of $1.67 \sqrt{f_{c}^{'}} MPa$ $(20 \sqrt{f_{c}^{'}} psi)$ or $1.25 \sqrt{f_{c}^{'}} MPa$ $(15 \sqrt{f_{c}^{'}} psi)$ for an interior joint with or without transverse beams, respectively, and $A_{j}$ is the effective cross-sectional area of the joint. Definitions of $A_{j}$ can be found in ACI 318 and ACI 352R-02. This article uses the one adopted in ACI 318-14, which is equal to the gross area of the column for the test specimens shown in Figure 3.

The strength reduction factor $ϕ$ is omitted for experimental studies and then the joint shear demand-to-capacity ratio is

\frac{V_{u}}{V_{n}} = \frac{V_{u}}{γ \sqrt{f_{c}^{'}} A_{j}} = \frac{v_{u}}{γ \sqrt{f_{c}^{'}}}

(5)

where $v_{u}$ is the design shear stress of the joint and equal to 10.9 MPa (A_j = 500 × 500 mm² and V_u = 2720 kN from equation (1)) for all test specimens.

Notably, the specified nominal shear strengths of specimens A1, A2, B1 and B2 (joint without transverse beams) is equal to $1.25 \sqrt{f_{c}^{'}} MPa$ but that of specimens A3 and B3 (joint with transverse beams) is $1.67 \sqrt{f_{c}^{'}} MPa$ . Due to set-up limitation, each transverse beam stub only had a length of 150 mm, which is less than one overall beam depth, and may not be effective for confinement. Therefore, all the bars in the transverse beams are anchored by heads to shorten the development length and promote the confinement effect.

Using the design concrete strength of 55 MPa, the joint shear demand-to-capacity ratio is equal to 1.18 for the test joints without transverse beams but equal to 0.88 for the test joints with transverse beams. In other words, the design joint shear stress in specimens A1, A2, B1 and B2 is 18% higher than the permissible value; thus, these joint specimens are expected to achieve beam yielding followed by joint shear failure at a limited inter-storey drift ratio. However, specimens A3 and B3 have a design joint shear stress below the permissible value and are expected to develop beam hinging until joint shear failure occurs at a relatively large drift ratio.

Provided-to-required transverse reinforcement ratio in the joint

To maintain the integrity of the joint concrete and to reduce the degradation of the joint shear capacity, transverse reinforcement required in the column ends shall be extended through the joint, unless a joint is considered to be effectively confined by beams on all four sides where the required amount of transverse reinforcement shall be permitted to be reduced by one-half, as given in Section 18.8.3 of ACI 318-14

\frac{A_{sh}}{s b_{c}} \geq 0.3 (\frac{A_{g}}{A_{ch}} - 1) \frac{f_{c}^{'}}{f_{yt}} and 0.09 \frac{f_{c}^{'}}{f_{yt}}

(6)

where $A_{sh}$ is the total cross-sectional area of the transverse reinforcement, including crossties, within spacing s and perpendicular to dimension $b_{c}$ , which is the cross-sectional dimension of the column core without concrete cover; $A_{g}$ and $A_{ch}$ are the gross area and the core area of the column, respectively; and $f_{yt}$ is the specified yield strength of the transverse reinforcement.

Kim and LaFave (2007) identified that an ‘ $A_{sh}$ ratio’ (provided amount of joint transverse reinforcement divided by the amount required by equation (6)) could be a key influencing parameter for the shear behaviour of beam–column joints. All the connections in this test programme were well confined by transverse reinforcement and had ‘ $A_{sh}$ ratio’ exceeding 1.0, as shown in Table 1.

Anchorage in beam–column joint

For Grade 420 straight beam bars passing through the joint, ACI 318 and ACI 352R-02 require a minimum column depth of 20 bar diameters $(20 d_{b})$ for joints of special moment frames. All test connections shown in Figure 3 used Grade 420 25-mm-diameter longitudinal reinforcing bars and a column depth of 500 mm, which is approximate to $20 d_{b}$ . For headed deformed bars anchored in the joints of special moment frames, ACI 318 gives a minimum development length, measured from the beam–column interface to the bearing face of the head, as follows

ℓ_{dt, 318} = the greatest of \frac{0.19 f_{y}}{\sqrt{f_{c}^{'}}} d_{b}, 8 d_{b}, 150 mm

(7)

with limitations of (a) specified $f_{y}$ not exceeding 420 MPa, (b) bar size not exceeding 36 mm, (c) normal-weight concrete, (d) each head has a net bearing area exceeding four times the nominal cross-sectional area of the bar, (e) minimum clear cover of $2 d_{b}$ for each bar and (f) minimum clear spacing of $3 d_{b}$ between parallel bars.

The recommendations of development length for headed bars given by ACI 352R-02 are somewhat different. For joints of special moment frames, critical sections for the development of beam bars should be taken at the outside edge of the confined core, and the minimum development length for headed bars is

ℓ_{dt, 352} = the greatest of \frac{0.15 f_{y}}{\sqrt{f_{c}^{'}}} d_{b}, 8 d_{b}, 150 mm

(8)

and bar heads should be located in the confined core within 50 mm from the back of the confined core (ACI-ASCE Committee 352, 2002).

Substituting bar $f_{y}$ of 420 MPa and design concrete strength $f_{c}^{'}$ of 55 MPa into equations (7) and (8) resulting in minimum development lengths of 10.8d_b and 8.5d_b, respectively. The latter should add a cover thickness of 25 mm (1d_b) for comparison with test specimens shown in Figure 3. The provided development length of the headed bars is only 9d_b, measured from the bearing face of the head to the column face. Furthermore, all test specimens were detailed with longitudinal beam bars at a clear spacing of 2d_b between bars. Obviously, the anchorage conditions of the bottom beam bars used in this test programme are relatively severe. Kang et al. (2009) extensively reviewed studies on the use of headed bars and concluded that the minimum clear spacing between headed bars can be reduced to 2d_b in beam–column joints confined by transverse reinforcement. Thereafter, Kang et al. (2012) tested two exterior beam–column joints with closely spaced headed bars and concluded that the clear bar spacing of approximately 2d_b or the use of two bar layers may be reasonably permitted for headed bars anchored in the exterior beam–column joints subjected to earthquake-type loading. Similar conclusions are also drawn by Chiu et al. (2016). Therefore, this experimental programme decided to use a minimum clear spacing of 2d_b between beam bars anchored in the joints.

Figure 4 shows the stress–strain relations obtained from the tensile tests of Grade 420 D25 bare and headed bars. The 65-mm-diameter steel head was attached to a reinforcing bar end by a friction–welding connection, which obstructed some of the bearing face of the head; thus, the net bearing area of the head was only about 4.3 times the nominal cross-sectional area of the bar. As shown in Figure 4, the bar-to-head connections are strong enough to break the bar in tensile tests resulting in stress–strain behaviour identical to that of bare bars.

Figure 4.

Stress–strain curves obtained from bar tensile tests (200-mm gauge length).

Test method and load protocol

Figure 5 shows the test set-up and reinforcing details in this experimental programme. The bottom column was hinged to the strong floor with a one-dimensional rocking steel base. The beams were laterally braced to restrain out-of-plane deformation and linked to the strong floor using roller supports at the right and left ends. At the beginning of each test, a constant axial load of $0.05 A_{g} f_{c}^{'}$ was applied to the column top by a hydraulic jack and pretension steel rods. The upper column was horizontally connected to the reaction wall by one displacement-controlled actuator to apply cyclic loading in a quasi-static manner. The test set-up was arranged to eliminate the P-delta effect and simulate a 3/4-scale beam–column joint specimen with a column height (storey height) of 3 m and a beam span of 4.5 m (the distance between the roller-supported inflection points).

Figure 5.

Overview of typical test set-up and details of cross section.

A typical displacement-controlled loading protocol consisting of three reversed cycles at gradually increased drift ratios (0.50%, 0.75%, 1.0%, 1.5%, 2%, 3%, 4%, 6% and 8%) was used in this study. The target displacement at the loading point of the upper column was computed by multiplying the target drift ratio to the simulated storey height of 3000 mm. The axial load and lateral force at the upper column were monitored by load cells. Several displacement transducers were attached to the test specimen to measure the global lateral drifts and local deformations (joint shear deformation and beam end rotation). Numerous strain gauges were pre-attached to reinforcements at key locations to record the strain histories. In general, the loading protocol and test procedure in this experimental programme are consistent with respect to ACI 374.1-05 (ACI Committee 374, 2005). The presented test results herein continued up to 6% or 8% drift ratio for the observation of failure modes. However, the performance of test specimens should be evaluated prior to the limiting drift ratio of 4% because the 6% or 8% drift may be too large for a well-designed special moment frame.

Experimental results and discussion

The experimental results showed that each test connection was capable of developing yielding in beams followed by joint shear failure (‘BJ’ failure) at a drift ratio of 4% or 6%, due to the difference in joint shear demand-to-capacity ratios. All beam longitudinal reinforcements developed the bar yield strength and underwent inelastic load reversals without bond or anchorage failure in the joint. The measured and observed responses are summarized and discussed in the following subsections.

Global response

Figure 6 shows the cyclic lateral load–displacement curves for all test specimens, where the drift ratio is equal to the lateral displacement $(δ)$ at the loading point divided by the simulated storey height of 3000 mm (Figure 5). The lateral load Q is also normalized to the ideal yield load (Q_y = 430 kN) in balance with the sum of the nominal flexural strengths of the beams calculated at the joint faces using the measured material properties (Table 1). Figure 6 also illustrates the displacement ductility ratio $μ = δ / δ_{y}$ in the positive and negative directions for each test specimen, where the idealized yield displacement $δ_{y}$ was equal to $Q_{y}$ divided by a secant stiffness obtained at the measured displacement corresponding to 0.75Q_y in the first cycle of 1.5% drift. The definition of $δ_{y}$ value can be found in ACI 374.2R-13 (ACI Committee 374, 2013).

Figure 6.

Lateral load–displacement responses for test specimens.

Typically detailed specimens A1 and B1, with continuous beam bars passing through the joint, developed yielding of beam bars in 1.5% drift cycles and attained maximum lateral resistance $(Q_{m})$ at 3% drift ratio, as shown in Figure 6, followed by strength and stiffness degradation due to shear cracks and damage in the joint. The maximum experimental shear stresses acting on the joint, back calculated using $Q_{m}$ and measured material properties, for specimens A1 and B1 were 11% and 9%, respectively, higher than the nominal strength of $1.25 \sqrt{f_{c}^{'}} MPa$ specified by ACI 318-14 for interior joint without transverse beams. Measured shear deformations on the joints of Specimen A1 and B1 exhibited inelastic hysteresis in 4% drift cycles followed by joint shear failure. Neither Specimen A1 nor B1 showed any bond failure along the straight beam bars in the joint during testing, and thus, the development length of 20d_b is concluded to be adequate for the test specimens.

As shown in Figures 6 and 7, the hysteresis and envelope curves of Specimen A2 with alternative reinforcing details for the anchorage of the bottom beam bars are almost identical to those of Specimen A1; however, those of Specimen B2 are somewhat inferior to those of Specimen B1. Although the global responses of the four cruciform beam–column connections were very similar during testing, a little more strength degradation can be observed in Specimen B2. This can be attributed to the numbers of bottom beam bars anchored in the middle of the joint. However, Specimens A3 and B3 with transverse beams obviously perform better than the other four cruciform specimens (Figure 6). Both specimens developed beam hinging, attained their maximum lateral resistance $(Q_{m})$ at a larger drift ratio of 6%, and then completed the 8% drift cycles for failure observation. The Q_m-back-calculated experimental shear stresses in the joints of Specimens A3 and B3 are 13% and 16% less than the nominal value of $1.67 \sqrt{f_{c}^{'}} MPa$ specified by ACI 318-14. Due to the presence of the transverse beams, the joint damage in Specimens A3 and B3 was not as severe as cruciform Specimens A2 and B2.

Figure 7.

Comparison of envelope curves for test specimens: (a) Group A specimens and (b) Group B specimens.

Crack observations

Figure 8 compares the crack patterns observed at the 3% drift ratio. In addition to the primary shear cracks that propagated diagonally in the joints, secondary cracks also initiated along the bottom beam bars anchored by heads in the joint, especially in Specimen B2. Increasing the number of bottom beam bars that terminate in the middle of the joint by heads seems to increase the potential for concrete breakout. As addressed in the commentary of Section 25.4.4.2 in ACI 318-14, breakout failure can be precluded in joints by either keeping the anchorage length from exceeding 0.66d, where d is the effective depth of the beam section, or by providing adequate transverse reinforcement to enable a strut-and-tie mechanism. The provided anchorage length of the headed bars in the joints of the test specimens is only about 0.5d, and thus, the role of the transverse reinforcement becomes crucial. Figure 9 illustrates a strut-and-tie model, where the flexure tension that developed in the bottom beam bars was anchored by heads and resisted by fan-shaped struts and transverse reinforcement. The anchorage performance of the bottom beam bars and the force-transferring behaviour of the transverse reinforcement can be monitored by strain gauge readings.

Figure 8.

Crack patterns observed at 3% drift ratio.

Figure 9.

Strut-and-tie modelling for Specimen A2 with headed bars anchored in the joint middle.

Local response

The strains of reinforcing bars were measured using electrical resistance strain gauges attached to some reinforcing bars at selected locations. Figure 10 shows the strain profiles along the bottom beam bars at peak drift ratios for the four cruciform specimens. All beam bar strains measured at the beam–column faces (location at ±250 mm) went above the ideal yield strain of 2155 µε (obtain from Figure 4) at the 2% drift ratio indicating the development of beam yielding. For specimens A1 and B1, a clear strain gradient per distance along the straight beam bar passing through the joint can be observed in Figure 10. This indicated that the bond resistance in the joint had not been completely destroyed till the 3% drift ratio. It is concluded that the development length of 20d_b is adequate for the straight beam bars used in the test specimens. For specimens A2 and B2 with alternative reinforcing details, the beam bar strains measured at the beam–column faces (location at ±250 mm) and in the joint core (location at ±100 mm) exceeded the bar yield strain at the 2% drift ratios indicating almost the entire bar tensile force was transferred to the end anchorage. It is evident that the bottom beam bars in specimens A2 and B2 were effectively anchored by heads with a short anchorage length and a clear spacing of 2d_b in the beam–column joints. Notably, the joints were well detailed with transverse reinforcement.

Figure 10.

Profiles of strain developed along bottom beam bars for test specimens.

Figure 11 compares the profiles of tie bar strains measured at one top column crosstie (Gauge 11) above the top beam bars, three inner joint crossties (Gauges 12–14) between the top and bottom beam bars and another bottom column crosstie (Gauge 15) below the bottom beam bars for the cruciform test specimens. Each gauge was attached to the centre of the crosstie parallel to the beam bars. Figure 11 shows that the tie bar strains measured in the joint (Gauges 12–14) remained elastic in the 1.5% drift cycles but went beyond the yield strain of 2500 µε at the 2% drift ratios for all test specimens. Notably, the tie bar strains measured in the top and bottom column (Gauges 11 and 15) remained elastic over the entire loading history in Specimens A1 and B1 with continuous bottom beam bars. In contrast, the tie bar strains measured below the discontinuous bottom beam bars in Specimens A2 and B2 (Gauge 15) went beyond the yield strain of 2500 µε after the 2% drift cycles. Besides, the tie bar strain of Gauge 14 in Specimen B2 was relatively larger than that of Gauge 14 in Specimen B1 for each drift level. These phenomena can be explained using Figure 9, where the tensile force of the bottom beam bars was transferred to the transverse reinforcement in the joint and the bottom column.

Figure 11.

Profiles of tie bar strain along the column height.

For the anchorage of the bottom beam bars by heads in the middle of the joint, yielding of the column transverse reinforcement is not preferred because it may affect the column confinement. Therefore, it is recommended to provide one more set of joint transverse reinforcements below the bottom beam bars anchored in the joint, as shown in Figure 12. To preclude breakout failure, this article recommends that the total amount of joint and column transverse reinforcements covered by the fan-shaped struts should be capable of resisting the total tensile force to be developed in the bottom beam bars, as given below

\sum A_{sh} f_{yt} \geq A_{s, bot} f_{y}

(9)

where $\sum A_{sh}$ is the amount of effective transverse reinforcement in the joint and bottom column, parallel to the bottom beam bars and within the critical edge distance of 1.5 times the effective embedded depth of the headed bars in the confined core, which is based on a breakout prism angle of approximately 35° (ACI 318-14 Chapter 17).

Figure 12.

Breakout failure precluded in joint by providing adequate transverse reinforcement to tie the potential failure surface.

For Specimens A2 and B2, the effective embedded depth of the headed beam bars in the confined column core was about 8d_b (200 mm), and thus, all three sets of joint transverse reinforcements and the upper two sets of transverse reinforcements in the bottom column were effective to tie the assumed breakout prism. Using equation (9) with a measured yield strength of $f_{yt} = 498 MPa$ and $f_{y} = 470 MPa$ , the total tie force provided by five sets of transverse reinforcements is 33% higher than the yield force of the four D25 bars in the beam bottom of Specimen A2 but 12% less than that of the six D25 bars in the beam bottom of Specimen B2. Therefore, with respect to Specimen A2, Specimen B2 exhibited relatively severe damage (Figure 8) and strength degradation (Figure 7).

Evaluation of seismic testing performance

Precast frame systems with alternative details that do not satisfy the prescriptive requirements of chapter 18 in ACI 318-14 may be used for seismic design if satisfactory performance can be demonstrated by experiments of the as-built design. ACI 374.1-05 defines a protocol for the design, analysis and laboratory testing of such frames. The proposed alternative reinforcing details for the emulative precast beam–column connections were designed and tested in this study according to ACI 374.1-05. For acceptance, the test results of the third complete cycle to a limiting drift ratio not less than 3.5% should satisfy the following three criteria:

Strength degradation shall not exceed 25% of the maximum peak strength in the same loading direction.

Secant stiffness between drift ratios of −1/10 and +1/10 of the limiting drift ratio shall not be less than 5% of the initial stiffness obtained from the first cycle.

Energy dissipation in the third cycle of limiting drift ratio shall not be less than 12.5% of the idealized elastoplastic energy of that drift ratio.

In this experimental programme, a limiting drift ratio of 4% is conservatively considered due to the lack of 3.5% drift cycles. Accordingly, Table 2 compares the hysteresis performance of the third cycle of the 4% drift ratio for the test specimens. All test specimens satisfied the acceptance criteria at a 4% drift ratio except that the strength degradation in the positive loading direction of Specimen B2 was close to 25% and, therefore, not ideal. This is attributed to the numbers of bottom beam bars in excess of the amount of effective transverse reinforcement in Specimen B2. On the contrary, Specimen A2 with alternate reinforcing details performed as well as and perhaps even better than the benchmark Specimen A1 with continuous bottom bars because of fewer bottom beam bars anchored in the joint. To conclude, the amount of effective transverse reinforcement should be proportioned to the pull-out of the headed bars from the joint. Based on the observation of the strain profiles of Figure 11, it is recommended to provide more transverse reinforcement adjacent to the bottom beam bars anchored in the joint by heads.

Table 2.

Test results for comparison with acceptance criteria of ACI 374.1-05.

Parameters		Specimen
Parameters		A1	A2	A3	B1	B2	B3
Initial stiffness for first cycle of 0.5% drift (kN/mm)	Positive	12.26	11.48	12.14	12.49	12.15	12.90
Initial stiffness for first cycle of 0.5% drift (kN/mm)	Negative	11.83	11.86	11.89	12.64	12.03	12.26
Percentage degradation in lateral resistance from first to third cycle of 4% drift	Positive	22.5%	18.8%	10.6%	18.6%	24.9%	10.6%
	Negative	16.2%	10.7%	2.8%	14.7%	20.8%	5.1%
Ratio of secant stiffness^a around zero drift to initial stiffness	Positive	0.13	0.15	0.19	0.15	0.11	0.16
	Negative	0.12	0.13	0.17	0.11	0.10	0.14
Relative energy dissipation ratio^b	E_D/E_PP	0.25	0.24	0.25	0.25	0.24	0.23

The secant stiffness around zero drift ratio was obtained for positive and negative loading directions between drift ratios of −0.4% and +0.4% in the third compete cycle of 4% drift ratio.

The relative energy dissipation ratio was the energy dissipated in the third complete cycle of 4% drift ratio divided by the idealized elastoplastic energy of that cycle.

Finally, as compared in Table 2, the performance of Specimens A3 and B3 with transverse beams is better than that of the other cruciform specimens. Obviously, the presence of the transverse beams did enhance the joint integrity even though the joint transverse reinforcement was also reduced. The transverse beams with discontinuous bottom beam bars framing to the joint may not be considered effective for confining the joint. This experimental programme is limited to emulative precast beam–column connections under unidirectional load reversals. It may be more crucial for such connections under bidirectional load reversals. This should be verified in future experimental studies.

Design recommendations

To ease the difficulty of the erection and fabrication of precast beams at cast-in-place beam–column connections, it is a viable option to terminate the protruded bottom bars of precast beams in the middle of the joint with adequate anchorage and confinement. On the basis of the experimental results and observations, the confinement and anchorage of the bottom beam bars in the joint is considered to be adequate if following conditions are satisfied:

Joint and column transverse reinforcements should conform to Section 18.7.5 of ACI 318-14.

The development length of the headed deformed bars should be in accordance with Section 25.4.4 of ACI 318-14, except that a minimum clear spacing of 2d_b between bars could be used in the confined joint.

Breakout failure of the headed bars should be precluded by providing adequate transverse reinforcement to enable a fan-shaped strut-and-tie mechanism, where the amount of transverse reinforcement within the critical edge distance (1.5 times the effective embedded depth of the headed bars in the confined core) should be proportional to the tensile force of the bottom beam bars.

For the test specimens, the clear spacing between beam bars was only 2d_b, and the provided development length of the headed bars was a little shorter than the required length given by ACI 318-14 or ACI 352R-02. Such headed bars did perform well in a well-confined beam–column joint but may not perform as well under other conditions. The test results show that the anchorage requirements of ACI 318-14 are conservative for a well-confined beam–column joint. In addition, it is unnecessary and not recommended to anchor top beam bars in the middle of beam–column joints, unless the top beam bars cannot be extended through the joint. As shown in Figures 1 and 2, straight and continuous top beam bars are preferred in cast-in-place joints and beam tops.

Conclusion

The experimental results presented in this work demonstrated that emulative precast beam–column connections with bottom beam bars anchored in the joint middle can perform as well as monolithic beam–column connections with straight beam bars passing through the joint. Based on the experimental results, it is concluded that the bottom beam bars protruding from precast beam units could be anchored in the joint middle by heads with limitations of a least net bearing area of $4 A_{b}$ for each head, a minimum clear spacing of 2d_b between bars, a code-conforming anchorage length and adequate transverse reinforcement.

From the experimental observations, the potential for concrete breakout failure increased for the closely spaced headed bars that were used in the test specimens. To preclude breakout failure, adequate transverse reinforcement should be provided and distributed uniformly within the critical edge distance of 1.5 times the effective embedded depth of the headed bars in the confined core. The amount of transverse reinforcement could be proportioned by establishing load paths in accordance with strut-and-tie modelling principles. Finally, with proper design and details, anchorage of bottom beam bars in the middle of the joint can be a viable option to relieve the steel congestion in the cast-in-place connections for special moment-resisting frames.

Footnotes

Declaration of Conflicting Interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship and/or publication of this article.

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship and/or publication of this article: This study was supported by the National Science Council (renamed to Ministry of Science and Technology since 2014) in Taiwan.

References

ACI-ASCE Committee 352 (2002) 352R-02: Recommendations for Design of Beam-Column Connections in Monolithic Reinforced Concrete Structures. Farmington Hills, MI: American Concrete Institute, p. 38.

ACI Committee 318 (2014) 318-14: Building Code Requirements for Structural Concrete and Commentary. Farmington Hills, MI: American Concrete Institute, p. 520.

ACI Committee 374 (2005) 374.1-05: Acceptance Criteria for Moment Frames Based on Structural Testing and Commentary. Farmington Hills, MI: American Concrete Institute, p. 9.

ACI Committee 374 (2013) 374.2R-13: Guide for Testing Reinforced Concrete Structural Elements under Slowly Applied Simulated Seismic Loads. Farmington Hills, MI: American Concrete Institute, p. 18.

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