Seismic behavior of encased CFT column base connections

Engineering Structures 182 (2019) 363–378

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Seismic behavior of encased CFT column base connections Xian Li

a,b,c,⁎

a

a

a,d

, Tao Zhou , Jian Li , Xiao-Bo Kuang

a,c

, Yu-Wei Zhao

T

a

Jiangsu Key Laboratory of Environmental Impact and Structural Safety in Civil Engineering, China University of Mining and Technology, Xuzhou 221116, Jiangsu, China State Key Laboratory for Geomechanics & Deep Underground Engineering, China University of Mining and Technology, Xuzhou 221116, Jiangsu, China Jiangsu Collaborative Innovation Center for Building Energy Saving and Construction Technology, Xuzhou 221116, Jiangsu, China d China Construction Engineering Company Limited, Beijing 100078, China b c

ARTICLE INFO

ABSTRACT

Keywords: Concrete filled steel tubes Column base Seismic behavior Connections Failure modes

The vulnerability of traditional encased column base connections, which consist of inner base-plate column bases partially encased by reinforced concrete outer components, has been demonstrated by previous earthquakes. This paper presents a new type of encased column base connections using steel wraps to confine reinforced concrete outer components, and the seismic behavior of both the traditional and the proposed encased concrete filled steel tube (CFT) column base connections was experimentally studied by testing of eight large-scale specimens under simulated seismic loads. The test parameters mainly included thickness, height and flexural reinforcement ratios of outer components, with or without steel wraps and shear studs. The effects of the test parameters on the failure modes, load-deformation response curves, stiffness, strength, ductility, energy dissipation ability and strain distribution of the encased column base connections were comprehensively evaluated. The test results indicate that encased column base connections with properly design can achieve desirable seismic behavior and the use of steel wraps can effectively prevent the shear cracking of reinforced concrete outer components. Finally, the methods of predicting the flexural strength of encased CFT column base connections were also discussed.

1. Introduction Concrete filled steel tubes (CFTs) have gained increasing popularity in industrial plants, high-rise buildings and bridge piers in recent years because of their structural and constructional merits. The favorable seismic performance of CFT structures can be guaranteed only when adequate design details are employed. Previous studies on CFT structures were primarily focused on the basic mechanical properties of CFT elements [1–3] and the seismic behavior of CFT column to beam/slab connections in buildings [4–7]. However, little attention was paid to the CFT column to footing connections. In a CFT structure, the loads such as gravity and seismic loads carried by columns must be transferred down through the footings to the soil/rock, and thus robust column base connections are quite important for reliable CFT construction. Existing types of CFT column base connections mainly include exposed base-plate connections, embedded connections and encased column base connections. The exposed base-plate connection, where a base plate is connected to the concrete footing using some anchor bolts embedded into the footing in advance, has some constructional

advantages and has been commonly used in low- and medium-rise structures [8,9]. However, the anchor bolts were found to be seriously damaged under shear and tension during previous earthquakes [10]. The embedded connection, in which the column is directly embedded into the concrete footing, is more effectively to guarantee adequate its stiffness and strength [11–13]. However, the footing in this connection type is usually required to have an un-economical thickness to fully develop the strength and ductility of CFT columns. The encased column base connection, in which the column with a base-plate is partially encased by a reinforced concrete outer component, can provide high strength and stiffness with moderate constructional complexity [14,15]. However, serious shear failure of the reinforced concrete outer components in this connection type was observed during previous earthquakes. Therefore, the encased column base connections are not encouraged to be applied in the buildings located in high seismicity regions in China though this connection type has some constructional and cost-effective merits. To widen the use of encased column base connections in seismicity regions, the seismic behavior of encased column base connections should be comprehensively studied and the shear failure of reinforced

⁎ Corresponding author at: Jiangsu Key Laboratory of Environmental Impact and Structural Safety in Civil Engineering, China University of Mining and Technology, Xuzhou 221116, Jiangsu, China. E-mail addresses: [email protected], [email protected] (X. Li).

https://doi.org/10.1016/j.engstruct.2018.12.076 Received 18 October 2018; Received in revised form 22 December 2018; Accepted 23 December 2018 0141-0296/ © 2018 Elsevier Ltd. All rights reserved.

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concrete outer components in encased column base connections should be properly avoided. However, to date the behavior of encased column base connections has not been well studied yet. Akiyama et al. [16] reported a series of experimental studies on the cyclic behavior of encased column base connections for W-shaped steel columns; while Nakashima carried out seismic tests of some square tubular steel column base connections repaired using concrete encasement [17]. As for the encased CFT column base connections, only the studies conducted by Wang et al. [14] and Xu et al. [15] were issued. Wang et al. [14] tested two encased column base connections for circular CFT columns under cyclic loads, and the contribution of the central steel cages in the CFT columns to seismic resistance was evaluated. Xu et al. [15] evaluated the effects of heights of reinforced concrete outer components, shear studs and axial load levels on the cyclic performance of encased column base connections for hexagonal CFT columns. During their tests, serious punching shear cracking and diagonal shear cracking of the reinforced concrete outer components were observed, and the shear cracking was found to have significant adverse effects on the ductility and energy dissipation capacities of tested specimens. Due to the lack of comprehensive studies, the design specifications for encased column base connections are still rather rough. The Chinese design specifications CECS-230 [18] and GB50936 [19] related to the design of encased CFT column base connections are not clear, while both AISC 341-10 [20] and European code EC4 [21] even do not contain any guidelines for encased column base connections. The required height and thickness of reinforced concrete outer components and the effects of shear studs are not clearly addressed. The load transfer mechanism of this connection type is also still unclear. In this study, a new type of encased column base connections using steel wraps to confine reinforced concrete outer components was proposed, and the seismic behavior of both traditional and proposed encased CFT column base connections was evaluated by testing of eight large-scale connection specimens under simulated seismic loads. The test parameters mainly included thickness, height and flexural reinforcement ratios of outer components, with or without steel wraps and shear studs. The effects of these parameters on the failure modes, load-deformation response curves, stiffness, strength, ductility, energy dissipation ability and strain distribution of encased column base connections were discussed.

the thickness of outer components Toc and the column diameter D. The letter “A” was followed by “rs” or “cs” to distinguish the different outer component types. Herein the letters ‘rs’ represent an outer component confined by rectangular steel wraps, while the letters ‘cs’ represent an outer component confined by circular steel wraps. According to the instruction, the four specimens labeled as CFTC-2.5-0.5, CFTC-2.5-0.5s, CFTC-2.5-0.5r and CFTC-1.8-0.5, respectively, had the reinforced concrete outer components without the confinement of steel wraps, while the other four specimens labeled as CFTC-1.8rs-0.5, CFTC-2.5rs-0.5, CFTC-2.5rs-0.4 and CFTC-2.5cs-0.5 were the proposed connection types with reinforced concrete outer components confined by steel wraps. The steel wraps in these specimens had a thickness of 6 mm and each of them was evenly divided into three parts along the height by two slits with a depth of 130 mm, aiming to provide confinement only. Note that the letter “s” at the end of CFTC-2.5-0.5s represents that the CFT column tube was welded with shear studs at the encased region, while the letter ‘r’ at the end of CFTC-2.5-0.5r represents that the outer component had a reduced flexural reinforcement ratio. According to the above instructions, specimen CFTC-2.5-0.5 had an outer component with a height of 683 mm (2.5D) and a thickness of 139 mm (0.5D), while specimen CFTC-1.8rs-0.5 had a reinforced concrete outer component confined by rectangular steel wraps and the outer component had a height of 500 mm (1.8D) and a thickness of 139 mm (0.5D). All of the concrete footing had a uniform dimension of 1250 mm × 1250 mm in plan and 500 mm in depth. Each concrete footing was longitudinally and transversely reinforced at both the bottom and the top with 18 mm-diameter HRB400 deformed bars (hotrolled ribbed bar Ф18 with a nominal yielding strength of 400 MPa) averagely spaced at 140 mm. The thickness of concrete cover is 50 mm. The CFT column was located at the center of the concrete footing and contained a steel tube with a 273 mm outside diameter D and a 7 mmthickness wall. The seamless steel tube was made of Chinese Q345B steel with a nominal yielding strength of 345 MPa, and had a base plate to expand its compression area under an axial load. The square base plate for all specimens had an identical dimension of 322 mm in width and 20 mm in thickness, and a hole with a diameter of 193 mm at the center. The square base plate was anchored to the concrete footing using four Grade 4.6 anchor bolts with a diameter of 20 mm at the corners. The Grade 4.6 anchor bolts have a nominal tensile strength of 400 MPa and a nominal yield strength of 240 MPa. The differences among these eight specimens were the details of their encased column base connections. Specimen CFTC-2.5-0.5 was designed as a benchmark. The reinforced concrete outer component was reinforced with twenty 16 mm diameter HRB400 flexural reinforcement and 8 mm diameter HPB300 stirrups with a spacing of 120 mm. The top of the outer component was heavily reinforced with four 8 mm diameter stirrups with a spacing of 50 mm. Specimen CFTC2.5-0.5s was designed to evaluate the effects of shear studs. In this specimen, forty-eight Chinese Grade 4.6 shear studs, with a 13 mm nominal diameter and a 50 mm-length, were welded around the CFT column base. The shear studs were uniformly distributed to six rows with eight studs for each row and a spacing of 100 mm between each two rows. To evaluate the effects of flexural reinforcement ratios of outer components, the outer component of specimen CFTC-2.5-0.5r was reinforced with twenty HRB400 deformed bars with a diameter of 14 mm. The outer component of specimens CFTC-1.8-0.5 and CFTC1.8rs-0.5 had a reduced height equal to 500 mm (about 1.8D). The reinforcement layouts of the outer components of specimens CFTC-1.8rs0.5 and CFTC-2.5rs-0.5 were the same to those of specimens CFTC-1.80.5 and CFTC-2.5-0.5, respectively. Compared to specimen CFTC-2.5rs0.5, specimen CFTC-2.5rs-0.4 had a reduced thickness of the outer component equal to 110 mm (about 0.4D). For specimen CFTC-2.5cs0.5, an outer component confined by circular steel wraps was employed and no stirrups were used in the circular outer component to simplify the construction. The detailed reinforcement layouts of specimens are shown in Fig. 1c–h.

2. Test program 2.1. Specimens design Eight large-scale specimens labeled as the form of ‘CFTC-A-B’ were constructed and tested, as shown in Table 1 and Fig. 1. The letters ‘CFTC’ represent concrete-filled steel tubular columns; the letter “A” represents the ratio of the height of outer components Hoc and the column diameter D, while the letter “B” represents the ratio between Table 1 Details of tested specimens. Specimens

Toc (mm)

Hoc (mm)

Flexural reinforcement

Shear stud

Steel wraps (b × t, mm)

CFTC-2.5-0.5 CFTC-2.5-0.5s CFTC-2.5-0.5r CFTC-1.8-0.5 CFTC-1.8rs-0.5 CFTC-2.5rs-0.5 CFTC-2.5rs-0.4 CFTC-2.5cs-0.5

139 139 139 139 139 139 110 139

683 683 683 500 500 683 683 683

20Ф16 20Ф16 20Ф14 20Ф16 20Ф16 20Ф16 20Ф16 20Ф16

wo w wo wo wo wo wo wo

– – – – □550 × 6 □550 × 6 □500 × 6 ○550 × 6

Note: w = with shear studs; wo = without shear studs; b = width or diameter; t = plate thickness. 364

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60

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150 130 110110

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36

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CFT column

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a) Dimension and reinforcement layouts of test specimens

38 85 80 96 80 85 38 500 A-A section

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g) CFTC-2.5rs-0.4

20 16 Anchor bolt

75 R2

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CFT column 273 7 Circular steel wrap 550 6

322 A-A section

h) CFTC-2.5rs-0.4

Fig. 1. Details of test specimens (Unit: mm).

All of the specimens were constructed by two steps as follows. Firstly, both the flexural reinforcement of outer components and the anchor bolts were properly placed in the reinforcing cages of footings and then the concrete of the footings was cast. About one week later, the steel tubes of CFT columns were anchored to the footings, and the formworks or the steel wraps of the outer components were erected. The infilled concrete of CFT columns and the concrete of outer components were poured subsequently.

2.2. Material properties The average yield stress fy and the tensile stress fu of steel components used in this test are summarized in Table 2. The concrete cubic strength fcu was determined by testing 150 mm concrete cubes at the time of specimen testing. The concrete was taken from the same batch of the specimens and was cured at room temperature the same to the test specimens. The average compressive strength of the concrete for footings was 52.2 MPa, while the corresponding value of infilled 365

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prescribed lateral drift ratios (Δ/H) included 0.25%, 0.5%, 0.75%, 1%, 1.5%, 2%, 3%, 4%, 5%, 6%, 7% 8%, 9% and 10%, where the drift ratio is defined as the applied displacement (Δ) at the shear force application point divided by the height (H = 1485 mm) from the loading point to the top surface of the concrete footing. For the drift ratios lower than 1%, one complete loading cycle was conducted for each drift ratio. For the drift ratios between 1% and 4%, each drift ratio was repeated three times to study the specimen’s strength degradation. The subsequent loading cycles higher than 4% were repeated twice and the test finished until some physical failure of the specimen was observed.

Table 2 Mechanical properties of steel components. Category

Call names

fy (MPa)

fu (MPa)

Steel tubes of CFT columns Steel wraps of outer components Stirrups Steel reinforcement Steel reinforcement Steel reinforcement

t = 7 mm t = 6 mm HPB300Ф8 HRB400Ф14 HRB400Ф16 HRB400Ф18

378 274 386 469 402 394

536 427 511 660 526 573

concrete in the columns and in the outer components was 63.4 MPa. The concrete cylinder strength f′c used in the following sections was taken as 0.8 times the cubic strength fcu.

3. Experimental results and discussions 3.1. Axial load distribution

2.3. Test methods

For the specimens at the elastic stage, the axial forces Nc resisted by CFT columns can be determined as follows according the measured strains ε.

The test setup for all specimens is shown in Fig. 2. During testing, the concrete footings of specimens were firstly tied to the rigid floor using eight high-strength bolts and four horizontal hydraulic jacks were used to prevent the potential slipping of the footing. The simulated seismic loads were applied to the CFT columns using a horizontally positioned 500 kN hydraulic actuator, while the axial load was applied to the top of the CFT columns using a cross beam post-tensioned by two vertical high-strength steel rods. The forces of the steel rods were applied by two 1000 kN capacity hydraulic hollow jacks and were measured by two calibrated load cells. To eliminate the bending of the rods during the lateral loading cycles, the bottom end of each rod was connected to the rigid floor using a specially designed pin device. The applied lateral displacement of the hydraulic actuator was monitored by both the displacement sensor itself and a separate linear potentiometer. The curvatures of the outer components at the bottom and the CFT columns above the outer components were monitored by linear potentiometers, as shown in Fig. 2. The corresponding lateral force was recorded by the load cell of the actuator. Electrical resistance strain gauges were affixed on the surfaces of steel tubes and steel bars at the sections A ∼ E shown in Fig. 3. The axial load distribution between CFT column bases and outer components was studied firstly, and the axial load was increased to 800 kN with an increment of 200 kN. During the subsequent quasi-static tests, the axial load was constantly maintained to 660 kN, and then displacement-controlled cyclic loads were applied to CFT columns. The

(1)

Nc = ·EA

where EA is the compression stiffness of CFT columns. Assuming that the axial load at Section E is totally resisted by the CFT column, the axial loads resisted by the CFT column at Sections B ∼ D can be determined according to the ratios between the strain εE at Section E and the strains εB, εC εD at Sections B ∼ D. The ratios K of the axial loads resisted by CFT columns at different sections are presented in Fig. 4. As shown in Fig. 4, the axial loads resisted by the encased CFT column bases decreased almost linearly with increase of the encased depth. For the specimens with 2.5D height outer components, the CFT columns in specimens CFTC-2.5-0.5, CFTC-2.5-0.5r and CFTC-2.5rs-0.5 carried about 50% of the total axial loads at Section D 610 mm above the footings. Due to the presence of shear studs in specimen CFTC-2.50.5s, the bond behavior between the CFT column and the outer component was improved and the CFT column in Section C carried a relatively low ratio of the total axial load. At Section B 50 mm above the footings, the axial loads of specimens CFTC-2.5-0.5s and CFTC-2.5rs-0.5 were almost totally carried by the outer components, while the outer components of specimens CFTC-2.5-0.5 and CFTC-2.5-0.5r carried about 80% of the total axial loads there. This result indicated that both the shear studs and steel wraps are efficient to improve the load transfer from CFT columns to outer components; however, the efficiency of the

Sensor Hydraulic jack

Reaction wall Loading beam

Actuator

LVDT5

Hydraulic jack

LVDT1

LVDT3

Loading plate High strength rod

LVDT4 LVDT2

Specimen

Steel beam Hydraulic jack

Fig. 2. Experimental setup. 366

B

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A

A

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a) Specimens with 2.5D height of outer components

b) Specimens with 1.8D height of outer components

Fig. 3. Cross sections of strain gauge arrangement.

steel wraps on the load transfer reduced owing to the reduction of outer component’s height. As shown in Fig. 4b, the axial distribution of specimens CFTC-1.8-0.5 and CFTC-1.8rs-0.5 was nearly similar. By calculating the areas enclosed by the curves and coordinate axes in Fig. 4, the ratios of the total axial loads resisted by CFT columns and outer components above Section B can be evaluated. For connections with 2.5D height outer components, the outer components carried about 75% of total axial loads, while the ratio of the axial load resisted by the 1.8D height outer components was about 60%. The shear studs and steel wraps had no significant effects on the total axial distribution above Section B.

was found at the top surface of the outer component during the drift ratio of 0.75% due to the punching of the CFT column. At the drift ratio of 1%, diagonal shear cracking occurred in the side surfaces of the outer component. At the drift ratio of 1.5%, the punching shear cracks with an angle of about 45 degrees to the loading direction were observed in the top surface of the outer component, and the separation of the CFT column from the outer component started. At the drift ratio of 3%, the diagonal shear cracking developed fast and some horizontal flexural cracks were observed around the section 150 mm above the footing. The width of the punching shear cracks at the top surface reached about 5 mm at the drift ratio of 4%. After that, the number and the width of the shear cracks increased rapidly as the applied displacements increased, and the specimen reached its peak load at the drift ratio of 6%. At the second cycle of 10% drift ratio, the concrete at the bottom section spalled off significantly and one anchor bolt fractured. No obvious buckling of the CFT column was observed during the testing. The failure appearance of specimen CFTC-2.5-0.5 at the final condition is presented in Fig. 5a. For specimen CFTC-2.5-0.5s, the separation between the CFT

3.2. Experimental observations and failure modes For the specimens with reinforced concrete outer components unconfined by steel wraps, their failure modes were primarily dominated by the serious diagonal shear cracking and punching shear cracking of the reinforced concrete outer components. During testing of the control specimen CFTC-2.5-0.5, the first cracking along the loading directions

1

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0 Section B

Section E

Section C

Section labels

Section D

Scetion labels

(a) Specimens with 2.5D height of outer components

(b) Specimens with 1.8D height of outer components

Fig. 4. Axial load distributions of the tested specimens. 367

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Concrete shear cracking

Concrete shear cracking

Steel tube fracture

Steel tube fracture

Stirrup exposure

Concrete crushing

(a) CFTC-2.5-0.5

Steel tube fracture

(b) CFTC-2.5-0.5s

(c) CFTC-2.5-0.5r

Slight buckling

Steel tube fracture

(d) CFTC-1.8-0.5

Slight buckling

Steel wrap fracture Concrete crushing

(e) CFTC-1.8rs-0.5

(f) CFTC-2.5rs-0.5

(g) CFTC-2.5rs-0.4

(h) CFTC-2.5cs-0.5

Fig. 5. Final failure appearances of the specimens after testing.

the failure appearances of the specimens after testing are presented in Fig. 5.

column and the outer component was effectively prevented due to the presence of shear studs. The failure of this specimen was dominated by the punching shear failure of the outer component above the shear studs and the final fracture of the CFT column. The diagonal shear cracking of the outer component was also found to be much less significant than that of specimen CFTC-2.5-0.5. For specimen CFTC-2.50.5r with a reduced flexural reinforcement ratio in the outer component, the specimen suffered more serious punching shear failure of the outer component than specimen CFTC-2.5-0.5 did. This may be attributed to the premature yielding of the flexural reinforcement and the subsequent serious separation between the CFT column and the outer component. The buckling of the CFT column at the final condition was not obvious in this specimen. For specimen CFTC-1.8-0.5, the punching shear failure of the outer component, serious concrete crushing at the bottom of the outer component and the fracture of CFT column above the outer component were observed at the final condition. The diagonal shear cracking in this specimen was much less significant than that of specimen CFTC-2.5-0.5. For the specimens with outer components confined by steel wraps, the number and the width of punching shear cracks at the top surfaces of outer components were significantly reduced due to the good confinement provided by the steel wraps, and the failure mode was mainly dominated by the plastic hinge failure of the CFT column above the outer components. The CFT columns in specimens CFTC-1.8rs-0.5 and CFTC-2.5rs-0.5 fractured at the drift ratios of 7% and 8%, respectively. The reduction of the outer component’s thickness in specimen CFTC2.5rs-0.4 resulted in a significant bulge of the steel wrap at the bottom and a fracture of one weld of the steel wrap finally occurred there. The CFT column in this specimen significantly buckled above the outer component but did not fracture until the test completed. At the final condition, the concrete at the bottom section was significantly crushed. In specimen CFTC-2.5cs-0.5, the outer component was confined by circular steel wraps and was not further reinforced by stirrups. The failure of this specimen was triggered by both the significant concrete crushing at the bottom section and the fracture of one flexural reinforcement there during a drift ratio of 9%. The steel tube also buckled at the region above the outer component. The detailed comparisons of

3.3. Load-deformation relationships The load-deformation relationship curves of specimens with unconfined reinforced concrete outer components are presented in Fig. 6a–d in terms of the applied shear loads versus the lateral displacements at the loading point. The hysteretic curves of the control specimen CFTC-2.5-0.5 showed some pinching due to the significant shear cracking of the outer component and the separation between the steel tube and the outer component. The specimen experienced a peak load of 357 kN and a maximum deformation of 10%. Specimen CFTC2.5-0.5r had a 3.1% reduction of the flexural reinforcement at each side of the outer component when compared with specimen CFTC-2.5-0.5. Consequently, the specimen showed 2.2% reduction of the loading capacity and more serious pinching of the hysteretic curves. Specimen CFTC-2.5-0.5s possessed improved bonding between the steel tube and the outer component due to the presence of shear studs. Thus, the specimen formed fatter hysteretic loops. The specimen showed an 8.3% enhancement of the loading capacity but a 29.3% reduction of the deformation capacity when compared with the control specimen CFTC2.5-0.5. These results can be explained by that the improvement of the bonding behavior between the steel tube and the reinforced concrete outer component provided a relatively good fixity at the top row of shear studs to the top pure CFT column. Therefore, the formation of a plastic hinge there led to a reduction of the distance between the load point and the plastic hinge. Specimen CFTC-1.8-0.5 exhibited stable spindle-type hysteretic loops with the most desirable deformation capacity (11.3%) among all of tested specimens since the significant diagonal shear cracking did not occur. That is to say, if the anchorage length of flexural reinforcements met to the requirements, the further increase of the heights of outer components would not improve the seismic behavior of the specimens significantly. Moreover, the testing of these specimens with unconfined reinforced concrete outer components can be found that the serious punching shear cracking and diagonal shear cracking should be properly prevented to obtain full hysteretic 368

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Lateral drift ratio (%) -8

-6

-4

-2

0

2

4

Lateral drift ratio (%) 6

8

10

12

-12 -10

500

500

400

400

300

300

200

200

Shaer load (kN)

Shear load (kN)

-12 -10

100 0 -100 -200

-6

-4

-2

0

2

4

6

8

10

12

0

30

60

90 120 150 180

100 0 -100 -200

-300

-300

-400

-400

-500

-8

-500 -180 -150 -120 -90 -60 -30

0

30

60

90 120 150 180

-180 -150 -120 -90 -60 -30

(mm)

(mm)

(g) CFTC-2.5rs-0.4

(h) CFTC-2.5cs-0.5 Fig. 6. (continued)

19.9 kN/mm in these four specimens, which was 28.4% higher than that of specimen CFTC-2.5-0.5. The reduction of the flexural reinforcement ratio and outer component’s height reduced the initial stiffness of the specimens. Specimens CFTC-2.5-0.5r and CFTC-1.8-0.5 had a comparable initial stiffness about 77.4% of the stiffness of specimen CFTC-2.5-0.5. However, the differences of the stiffness of these specimens became smaller with increase of the applied displacements. The peak-to-peak stiffness of the specimens with reinforced concrete outer components confined by steel wraps is plotted in Fig. 7b. Among these four specimens, specimen CFTC-2.5rs-0.5 had the highest initial stiffness; however, specimens CFTC-2.5rs-0.5 and CFTC-2.5cs-0.5 had nearly the same peak-to-peak stiffness when the applied drift ratios were larger than 1.5%. The reduction of the height and the thickness of outer components resulted in the reduction of the stiffness of the specimens. The further comparison of the stiffness of typical specimens with or without confined outer components is presented in Fig. 7c. The comparison of the initial stiffness of specimens with the same dimensions of the outer components can be found that the confinement of the outer components using steel wraps significantly increased the initial stiffness of specimens by about 30%. It is also interesting to find that specimens CFTC-2.5-0.5 and CFTC-1.8rs-0.5 with different heights of the outer components exhibited nearly the same peak-to-peak stiffness at each drift ratio due to the effect of steel wraps.

loops. The load-deformation relationship curves of the specimens with reinforced concrete outer components confined by steel wraps are presented in Fig. 6(e)–(h). Since these specimens failed due to serious plastic deformation of the steel tubes above the outer components, the load-deformation relationship curves of these specimens resembled those of pure concrete filled steel tube columns. Since the formation of plastic hinges there reduced the distance between the center of plastic hinge and the load point, all these specimens exhibited higher loading capacity but lower deformation capacity when compared with the specimens with unconfined reinforced concrete outer components. The comparisons between test results from specimens CFTC-1.8-0.5 and CFTC-1.8rs-0.5 indicated that steel wraps used in specimen CFTC-1.8rs0.5 resulted in 8.4% increase of the loading capacity but 32.2% reduction of the deformation capacity. It should be noted that, if specimen CFTC-1.8rs-0.5 was assumed to be fixed at the top surface of the outer component, both specimens CFTC-1.8-0.5 and CFTC-1.8rs-0.5 achieved comparable drift ratio capacities prior to failure. The comparisons between the test results from specimens CFTC-2.5rs-0.5 and CFTC-1.8rs-0.5 indicated that the increase of the height of the outer component had little effect on the hysteretic behavior of the specimens except the change of fixed points of the CFT columns. The reduction of the thickness of the outer component in specimen CFTC-2.5rs-0.4 resulted in obvious pinching of the hysteretic curves. Compared with specimen CFTC-2.5rs-0.5, specimen CFTC-2.5rs-0.4 was found to have a relatively low loading capacity but 11.1% increase of the deformation capacity due to the plastic deformation at the bottom section. The comparison of the test results of specimens CFTC-2.5rs-0.5 and CFTC2.5cs-0.5 indicated that the use of circular steel wraps improved both the loading and the deformation capacities of the specimen due to the good confinement provided by circular steel wraps.

3.5. Ductility The ductility of the tested specimens is presented in Table 3, where the displacement ductility factor μ is defined as the ratio between the ultimate displacement Δu and the yield displacement Δy. The ultimate displacement is obtained at 85% of the peak capacity at the descending stage of load-deformation envelop curve, while the yield displacement is determined by the tangent method. If some physical failure (such as steel tube fracture, anchor bolt fracture, reinforcement fracture or punching shear failure of the outer component, etc.) occurred prior to the gradual decrease of the loading capacity by 15%, the ultimate displacement will be adopted as the former last loading drift. As shown in Table 3, all of the tested specimens exhibited good ductility with the ductility factors varying between 4.30 and 7.19. The ductility factor of the control specimen CFTC-2.5-0.5 was 6.91. In general, the specimens that developed a plastic failure of CFT columns above the outer components exhibited a relatively reduced ductility. As explained previously, the formation of the plastic hinge above the outer

3.4. Stiffness and stiffness degradation The peak-to-peak stiffness presented in Fig. 7 was adopted herein to evaluate the stiffness characteristics of the tested specimens. The peakto-peak stiffness is defined as the average peak load at each deformation level divided by its corresponding displacement. The comparison of the initial stiffness of the specimens with un-confined reinforced concrete outer components was shown in Fig. 7a. The control specimen CFTC2.5-0.5 had an initial stiffness of 15.5 kN/mm. The specimen CFTC-2.50.5s with shear studs exhibited the highest initial stiffness equal to 370

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20

20

CFTC-2.5-0.5

16

Peak-to-peak stiffniss Ki(kN/mm)

Peak-to-peak stiffness Ki(kN/mm)

18

CFTC-2.5-0.5s

14

CFTC-2.5-0.5r

12

CFTC-1.8-0.5

10 8 6 4

CFTC-1.8rs-0.5

16

CFTC-2.5rs-0.5

14

CFTC-2.5rs-0.4

12

CFTC-2.5cs-0.5

10 8 6 4 2

2

0

0 0

2

4

6

8

10

12

Lateral drift ratio (%)

20

0

2

4

6

8

10

12

Lateral drift ratio (%)

(b) Specimens with confined outer components

(a) Specimens without unconfined outer components

Peak-to-peak stiffniss Ki(kN/mm)

18

CFTC-2.5-0.5

18 16

CFTC-1.8-0.5

14

CFTC-1.8rs-0.5

12

CFTC-2.5rs-0.5

10 8 6 4 2 0

0

2

4 6 8 Lateral drift ratio (%)

10

12

(c) Comparisons of stiffness of specimens with or without confined outer components Fig. 7. Stiffness degradation of tested specimens.

component reduced the distance between the load point and the center of plastic hinge, and thus reduced the absolute ultimate deformation of the specimen at the ultimate stage. Note that in practice the encasement length had less ratio of the column length than this test did, and thus the failure at the CFT section is a desirable failure mode. The test result also indicated that the use of shear studs in specimen CFTC-2.5-0.5s significantly reduced the ductility factor to 4.43, while specimen CFTC2.5-0.5r with a reduced flexural reinforcement ratio of the outer component exhibited the best ductility among these eight specimens. This is because the reduction of the flexural reinforcement ratio resulted in the corresponding reduction of the yield displacement of the specimen.

With sufficient stirrups, the loading capacity of specimen CFTC-2.5-0.5r deteriorated gradually and thus the specimen had a large ultimate deformation equal to 133.7 mm. Specimen CFTC-1.8-0.5 with a reduced height of the outer component showed a comparable ductility to that of specimen CFTC-2.5-0.5. Specimens CFTC-2.5rs-0.5 and CFTC-1.8rs-0.5 exhibited the ductility inferior to their counterparts with unconfined outer components due to the change of the locations of the plastic hinges. The comparison of the ductility factors of specimens CFTC2.5rs-0.5 and CFTC-2.5rs-0.4 indicated that the reduction of the thickness of the outer component confined by steel wraps increased the ultimate displacement of the specimen and thus slightly increased the

Table 3 Ductility factors of tested specimens. Specimens

Yield displacement (mm)

Yield load (kN)

Peak displacement (mm)

Peak load (kN)

Ultimate displacement (mm)

Ultimate load (kN)

Ductility factor μ

CFTC-2.5-0.5s CFTC-2.5-0.5r CFTC-1.8-0.5 CFTC-1.8rs-0.5 CFTC-2.5rs-0.5 CFTC-2.5rs-0.4 CFTC-2.5cs-0.5

23.5 18.6 24.3 24.2 23.5 24.3 27.6

335.2 270.0 250.5 276.5 323.2 330.5 370.4

74.3 74.3 74.3 66.8 74.3 81.7 74.3

386.5 349.0 315.5 342.0 423.5 402.0 438.0

104.0 133.7 163.4 104.0 104.0 142.7 118.8

352.0 318.0 275.5 310.5 380.5 341.7 423.0

4.43 7.19 6.72 4.30 4.43 5.87 4.30

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2.4

2.4

2

2

1.6

1.6

1.2

1.2

Eh

Eh

X. Li et al.

0.8

0.8

CFTC-2.5-0.5

CFTC-1.8rs-0.5 CFTC-2.5rs-0.5

CFTC-2.5-0.5s 0.4

0.4

CFTC-2.5-0.5r

CFTC-2.5rs-0.4

CFTC-1.8-0.5 0

0

a

2

4

6

8

10

CFTC-2.5cs-0.5 0

12

Lateral drift ratio(%)

0

2

b

Specimens without unconfined outer components

4

6 8 Lateral drift ratio(%)

10

12

Specimens with confined outer components

2.4 2

Eh

1.6 1.2 CFTC-2.5-0.5

0.8

CFTC-1.8-0.5 CFTC-1.8rs-0.5

0.4

CFTC-2.5rs-0.5 0

c

0

2

4

6

8

10

12

Lateral drift ratio(%)

Comparisons of stiffness of specimens with or without confined outer components Fig. 8. Energy coefficient curves of tested specimens.

ductility of the specimen. Specimen CFTC-2.5cs-0.5, which had a circular outer component confined by steel wraps, showed an improved deformation capacity when compared with specimen CFTC-2.5rs-0.5; however, specimen CFTC-2.5cs-0.5 had a slightly low ductility due to the increase of the yield displacement of the specimen.

at the same drift ratios the energy dissipation index slightly deteriorated as the number of loading cycles increased. Fig. 8a gives the energy dissipation index of specimens with unconfined outer components. Specimen CFTC-2.5-0.5s with shear studs had the highest energy dissipation index reaching about 2.2 at the drift ratio of 8%, while specimen CFTC-2.5-0.5r had the smallest energy dissipation index of about 1.1 at the drift ratio of 10% due to the significant pinching of the hysteretic loops. Specimens CFTC-2.5-0.5 and CFTC-1.8-0.5 had the energy dissipation index between those of specimens CFTC-2.5-0.5r and CFTC-2.5-0.5s. Fig. 8b gives the energy dissipation index of specimens with outer components confined by steel wraps. It can be found that both the reduction of the outer component’s thickness and the use of circular outer component significantly reduced the energy dissipation index of specimens. This is because the reduction of the thickness of outer component resulted in the premature crushing of concrete at the bottom of outer component while the plastic deformation of specimen CFTC-2.5cs-0.5 mainly concentrated into the section between the outer component and the concrete footing. The further comparison of the energy dissipation index of the typical specimens with or without confined outer components (CFTC-2.5-

3.6. Energy dissipation capacity The energy dissipation capacity of the tested specimens under cyclic loads was evaluated by both energy dissipation index Eh and cumulative dissipated energy, as shown in Fig. 8. The detailed definition of both the energy dissipation index Eh and the cumulative dissipated energy can be found in Li et al. [22]. A specimen with a larger energy dissipation index or a larger cumulative dissipated energy generally has greater energy dissipation capacity by plastic deformation. As shown in Fig. 8, during the elastic stage with drift ratios lower than 1%, all specimens had nearly the same and small energy dissipation index. The energy dissipation index kept almost constant at the stage from 1% to 2% drift ratios. After that, the energy dissipation index increased nonlinearly as the applied displacement increased. However, 372

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Cumulative energy dissipation (kJ)

800

2%. At subsequent large drift ratios, the strain kept almost about 2000με. The reason might be that the serious damage of the concrete in the outer component occurred at the large drift ratios and thus the contribution of the stirrups to shear resistance could not be effectively motivated. The stirrups at the bottom section of specimen CFTC-2.5-0.5 did not yield even at the final condition. For specimen CFTC-2.5-0.5s, the use of shear studs significantly increased strain demands of the stirrups at the top section. The strain of stirrups at the top section reached 11570με at the drift ratio of 7%. However, the stirrups at the bottom section showed a very small strain during the entire testing. For specimen CFTC-2.5-0.5r with a reduced flexural reinforcement ratio, the stirrups at the top section yielded at the drift ratio of 2%. As the applied displacement increased, the strain values of the stirrups at the top section did not further increase due to the serious shear cracking of concrete; however, the strain of the stirrups at the bottom section increased rapidly due to the pry-out actions there. For specimen CFTC1.8-0.5 with a reduced height of the outer component, the strains of stirrups at the bottom section slightly increased when compared with that of specimen CFTC-2.5-0.5. For the specimens with outer components confined by steel wraps, it is interesting that the strains of stirrups in the positive and negative loading directions were not symmetrical due to the presence of the steel wraps. Under the shear loads, the concrete breakout at the top of the outer component and the concrete pry-out at the bottom of the outer component may occur. As a result, relatively high strain in the stirrups far away the potential failure blocks was observed. It also can be found from Fig. 11 that the increase of the height and the thickness of the outer components resulted in a relatively rapid increase of the strain of stirrups, and the strains of the stirrups at the top section D increased more rapidly than that at the bottom section B. The stirrups at the bottom section B of all specimens did not yield even at the failure conditions of the specimens. 2) Strain distribution of steel wraps The typical transverse strain distributions of steel wraps along the column axis are presented in Fig. 12. As shown in the figure, the transverse strain was nearly linearly distributed along the height of the outer components. For specimen CFTC-1.8rs-0.5, the steel wraps at the section D yielded at the drift ratio of about 2%. The increase of the height of outer component reduced the strain demands of steel wraps, as evidenced by the fact that the steel wraps used in specimen CFTC2.5rs-0.5 yielded at the drift ratio of 5%. The comparison of the test results of specimens CFTC-2.5rs-0.5 and CFTC-2.5rs-0.4 indicated that the reduction of the thickness of outer component resulted in the increase of the strain of the steel wraps. At the drift ratio of 9%, the transverse strain at the section B of specimen CFTC-2.5rs-0.4 increased rapidly due to the serious concrete crushing there. 3) Strain distribution of the flexural reinforcement of outer components Fig. 13 shows the strain distribution of the outer components’ flexural reinforcement along the distance H to the top surface of the concrete footing. As shown in this figure, the flexural reinforcement of all specimens except specimen CFTC-1.8-0.5 yielded before the failure of specimens. The flexural reinforcement of the outer component of specimen CFTC-2.5-0.5 yielded at the bottom section B at the drift ratio of 5%, and then the strain firstly increased and then reduced as the applied displacement increased. The reason may be that at the large drift ratios the concrete significantly cracked and the bonding between the flexural reinforcement and the cracked concrete gradually deteriorated. Due to the use of shear studs in specimen CFTC-2.5-0.5s, the bond behavior between the CFT column and the outer component was improved and less moment was resisted by the outer component. Therefore, the measured strain of the flexural reinforcement was relatively small compared with that of specimen CFTC-2.5-0.5. The reduced flexural reinforcement ratio in specimen CFTC-2.5-0.5r resulted in the yield of the reinforcement at the drift ratio of 3%. In specimen CFTC1.8-0.5, the flexural reinforcement of the outer component did not yield

CFTC-2.5-0.5

700

CFTC-2.5-0.5s

600

CFTC-2.5-0.5r

500

CFTC-1.8rs-0.5

CFTC-1.8-0.5 CFTC-2.5rs-0.5

400

CFTC-2.5rs-0.4

300

CFTC-2.5cs-0.5

200 100 0

0

2

4

6

Drift ratio (%)

8

10

12

Fig. 9. Cumulative energy dissipation of tested specimens.

0.5, CFTC-2.5rs-0.5, CFTC-1.8-0.5 and CFTC-1.8rs-0.5) is presented in Fig. 8b. As shown in the figure, the energy dissipation index of specimen CFTC-1.8rs-0.5 and specimen CFTC-2.5rs-0.5 at the drift ratio of 7% were improved by about 20% and 30%, respectively, when compared with the corresponding values of specimens CFTC-1.8-0.5 and CFTC-2.5-0.5. This is mainly attributed to that the steel wraps effectively prohibited the shear cracking of outer components. The test results also indicated that the increase of the height of outer component had adverse effects on the energy dissipation of specimens. Thus in practical design the height of outer component is adequate when it can efficiently anchor the flexural reinforcement. The comparison of the cumulative energy dissipation of all tested specimens is presented in Fig. 9. The cumulative energy dissipation of the specimens increased rapidly after the drift ratio of 2%. The specimens with confined outer components generally showed a good energy dissipation capacity when the drift ratios were lower than 6%. At the drift ratio of 6%, the cumulative energy dissipation of specimen CFTC2.5rs-0.5 was 329.3 kJ, which was 30.6% higher than that of specimen CFTC-2.5-0.5. However, the specimens with unconfined outer components generally had a larger deformation capacity and thus dissipated a larger amount of cumulative energy before failure. Specimens CFTC2.5-0.5 and CFTC-1.8-0.5 dissipated a desirable cumulative energy equal to 781.9 kJ and 739.2 kJ at the drift ratio of 10%, respectively. The desirable energy dissipation of these two specimens was attributed to that the separation between CFT columns and outer components gradually increased plastic zones of columns and the large deformation of flexural reinforcement in outer components also contributed to the energy dissipation. 3.7. Strain analysis In these experiments, strain gauges were affixed onto the surfaces of the flexural steel reinforcement and stirrups of outer components as well as steel wraps to monitor their strain responses and strain distribution. According to the measured material properties, the yield strains of steel tubes for CFT columns and steel wraps were 1835με and 1330με, respectively, while the yield strains of the flexural reinforcement Φ14, Φ16 and stirrups Φ8 were 2345με, 2010με and 1930με, respectively. 1) Strain responses of stirrups of outer components Figs. 10 and 11 present the strain responses of the stirrups at the top and bottom sections of the outer components. For the specimens with un-confined reinforced concrete outer components, the strain-drift ratio relationship curves were symmetrical. The measured stirrup at the top section of specimen CFTC-2.5-0.5 yielded at the drift ratio of about 1.0% and got a high strain value equal to 4790με at the drift ratio of 373

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2000

14000 12000

1500

8000

Strain ( )

Strain ( )

10000 6000 4000 2000

1000 500 0

0 -2000

-500 -12

-8

-4

0

4

8

12

-12

-8

CFTC-2.5-0.5 CFTC-2.5-0.5r

-4

0

4

8

12

Lateral drift ratio (%)

Lateral drift ratio (%) CFTC-2.5-0.5s CFTC-1.8-0.5

CFTC-2.5-0.5 CFTC-2.5-0.5r

CFTC-2.5-0.5s CFTC-1.8-0.5

Fig. 10. Strain response of stirrups of specimen with unconfined outer components.

even at the final condition, indicating that the outer component resisted less part of the total moment due to the reduction of the outer component’s height. The comparison of the strain distributions of specimens with or with confined outer components indicated the use of steel wraps increased the strain of the flexural reinforcement since the steel wraps effectively inhibited the shear cracking of the concrete.

loading point, equal to 1485 mm herein; Mc is the flexural capacity of reinforced concrete outer component at the bottom section; Mu3 is the flexural capacity of the inner base plate connection. For the specimens having traditional outer components unconfined by steel wraps and without shear studs, lr is taken as the distance between the topmost stirrup of the outer component and the inner base plate. For other specimens having outer components confined by steel wraps or with shear studs, lr in Eq. (3) is proposed to be replaced by lr + D/2 due to the formation of plastic hinges there. According to Eqs. (2)–(4), if Mu2 ≥ Mu1, the encased column base connection develops a plastic hinge failure at the CFT section above the outer component. If Mu2 < Mu1, the failure at the outer component’s bottom section occurs.

4. Loading capacity analysis Since the flexural yielding of the tested specimens occurred prior to shear failure, the calculation of the shear strength was not discussed herein. For encased column base connections under flexure and axial loads, the failure at the CFT section above the outer component or at the outer component’s bottom section may occur. Thus, the flexural capacity of an encased column base connection Mu is determined by these two critical sections according to the following equations.

Mu = min(Mu1, Mu2)

(2)

Mu1 = Mpc /(1

(3)

lr /l)

4.1. Flexural capacity of CFT column Mpc The CFT section above the outer component is subjected to combined axial loads and flexural moment. In this study, the flexural capacity of CFT column was calculated according to EC4 code [21]. The steel tube was assumed to develop its full yield strength in the tension and compression zones, while the compressive strength of the confined concrete was taken as fc′. The predicted flexural capacity of CFT column Mpc is presented in Table 4.

(4)

Mu2 = Mc + Mu3

where Mpc is the full plastic bending capacity of an axially loaded CFT column; l is the distance between the steel column base plate and the

4000

1200

Strain gauge

3500

Strain ( )

Strain ( )

800

2500 2000 1500 1000

600 400 200

500

0

0 -500

Strain gauge

1000

3000

-12

-200

-8

-4 0 4 8 12 Lateral drift ratio (%) CFTC-1.8rs-0.5 CFTC-2.5rs-0.5 CFTC-2.5rs-0.4

-12

-8

-4 0 4 8 Lateral drift ratio (%) CFTC-1.8rs-0.5 CFTC-2.5rs-0.5 CFTC-2.5rs-0.4

Fig. 11. Strain response of stirrups of specimen with confined outer components. 374

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700 600

400 300

-0.5% -1% -1.5% -2% -3% -4% -5% -6%

200 100 0 -500

Distance to the footing (mm)

Distance to the footing (mm)

500

0

500

1000 1500 2000 2500 3000

500 400

-0.5% -1% -1.5% -2% -3% -4% -5% -6%

300 200 100 0 -500

0

Strain ( ) a) Specimen CFTC-1.8rs-0.5

500

1000

1500

2000

Strain ( ) b) Specimen CFTC-2.5rs-0.5

Distance to the footing (mm)

700 600 500

-0.5% -1% -2% -3% -4% -6% -8% -9% -10%

400 300 200 100 0 -1000

0

1000

2000

3000

4000

5000

6000

Strain ( ) c) Specimen CFTC-2.5rs-0.4 Fig. 12. Transverse strain distribution of steel wraps.

4.2. Flexural capacity of encased column base connection at the bottom section Mu2

is resisted by the concrete; 2) The effect of local compression under the base plate is neglected. The concrete stress of concrete is taken as 0.85fc′. The flexural strength of inner base plate connection Mu3 was calculated according to the strength model shown in Fig. 14 and the predicted values are presented in Table 4. The comparison between the predicted strength Mu and the measured yield strength Mty is presented in Table 4. As shown in Table 4, the proposed method gave a reasonable prediction of the tested specimens’ flexural strength with the ratios Mty/Mu varying between 0.98 and 1.11. The predicted results coincided well with the experimental observations that the yielding of CFT columns above the outer components occurred prior to the failure of outer components. The final failure of the outer components at the bottom occurred in some specimens due to the hardening of steel tubes at the large drift ratios.

In the study conducted by Xu et al. [15], the inner base plate connection was assumed to be subjected to combined axial compression and flexure at the connection’s bottom section while the outer component there was assumed to be subjected to bending only. The comparison of the analytical and test results indicated that the assumptions proposed by Xu et al. [15] resulted in too conservative prediction of the strength of encased column base connections. In this study, both the inner base plate connection and the outer component were assumed to be subjected to combined axial compression and flexure. According to the above test, the outer components with heights of 2.5D and 1.8D were assumed to resist 75% and 60% of the applied axial loads, respectively. The steel wraps for outer components are assumed to resist shear force only. In calculation of the flexural capacity of the outer component, the outer component was assumed to have an I-shaped cross section, which had a depth of the web equal to the diameter of the CFT column. According to the ACI code [23], the flexural capacity of the outer component Mc was calculated and listed in Table 4. In calculation of the flexural capacity of inner base plate connection Mu3, the following assumptions are made: 1) Under the moment, the tensile load is resisted by the anchor bolts, while the axial compression

5. Conclusions According to the testing of eight specimens under cyclic loads, some important conclusions were drawn as follows. (1) The encased column base connection has desirable seismic behavior and thus is a viable connection type for CFT structures in seismicity regions. The connection specimens had three typical final failure modes, including plastic hinge failure at the CFT column section or 375

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700

700

600

2%

500

5%

300

6%

200

7%

100

8%

1000

2000

3000

300

3%

200

5%

-100

4000

-1000

6% 7% 8%

0

1000

2000

3000

(b) CFTC-2.5-0.5s

(a) CFTC-2.5-0.5 700

500

600

400 0.5%

500

1%

400

3%

300

5%

200

6%

2000

3000

4% 6% 7% 8%

0

0

1000

3%

200 100

10%

0

2%

5%

8%

100

1%

300 H (mm)

H (mm)

4%

Strain ( )

Strain ( )

-100 -1000

2%

0

10%

0

1%

100

9%

0

0.5%

400

4%

H (mm)

H (mm)

500

3%

400

-100 -1000

600

1%

-100 -2000

4000

-1000

Strain ( )

0

1000

2000

3000

Strain ( )

(c) CFTC-2.5-0.5r

(d) CFTC-1.8-0.5 Fig. 13. Strain distribution of flexural reinforcement.

at the outer component’s bottom section, as well as the shear cracking failure of the reinforced concrete outer component. For the specimens failed at the CFT column section, the local buckling and final fracture of the steel tubes above the outer components occurred. For the specimens failed at the outer component’s bottom section, the concrete there significantly crushed with significant plastic deformation of the flexural reinforcement or fracture of some anchor bolts. The CFT columns also buckled to varying degrees above the outer components. For the specimens failed in shear cracking, the outer components suffered serious punching shear cracking and diagonal shear cracking. Significant separation between the CFT columns and the reinforced concrete outer components also occurred under the cyclic loads. (2) The use of steel wraps effectively prevented the shear cracking failure of the reinforced concrete components, resulting in plastic hinge failure at the CFT column sections or at the bottom sections of

the outer components. The specimens with outer components confined by steel wraps exhibited desirable hysteretic behavior, higher initial stiffness and better loading capacity. However, since the deformation of the outer components was prohibited, the deformation capacity and cumulative energy-dissipation capacity at the failure conditions of the specimens deteriorated. (3) The use of shear studs significantly improved the bond behavior between the CFT column and the outer component of the specimen. The specimen failed with less serious shear cracking of the outer component and final fracture of the CFT column above the outer component. Compared with the test results of the control specimen, the loading capacity and the initial stiffness of the specimen increased by 8.3% and 28.4%, respectively; however, the deformation capacity and the ductility decreased by 30.0% and 35.9% due to the formation of plastic hinge at the CFT section above the outer component.

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700

700

600

600

500

0.5% 1% 2%

300

3%

200

4%

H (mm)

H (mm)

400

5%

100

500

1%

400

2%

300

3%

200

4% 5%

100

6%

0 -100 -1000

0.5%

6%

0

0

1000

2000

3000

4000

-100 -1000

5000

0

Strain ( ) (e) CFTC-1.8rs-0.5

3000

4000

5000

Strain ( ) (f) CFTC-2.5rs-0.5

700

700

600

600 0.5%

500 400

1%

H(mm)

2%

300

3% 4%

200

400

2%

300

3% 4%

200

5%

100

0.5%

500

1%

H (mm)

1000 2000

5%

100

6%

0

6% 7%

0

-100 -2000 -1000

0

1000

2000

3000

-100 -1000

4000

Strain ( ) (g) CFTC-2.5rs-0.4

1000

3000

5000

7000

Strain ( ) (h) CFTC-2.5cs-0.5 Fig. 13. (continued)

(4) The reduction of the outer component’s height had adverse effects on the loading capacity and the initial stiffness of an encased column base connection but resulted in improved deformation capacity, ductility and energy dissipation capacity. The reduction of the flexural reinforcement ratio and the thickness of the outer

component led to notable pinching of the hysteretic loops with smaller energy dissipation index. (5) The methods for predicting the flexural strength of encased column base connections were proposed and the proposed methods could give reasonable results with good agreement with test results.

Table 4 Comparison of predicted and experimental flexural capacities. Specimens

Mpc (kN.m)

Mu1 (kN.m)

Mu3 (kN.m)

Mc (kN.m)

Mu2 (kN.m)

Mu (kN.m)

Mty (kN.m)

Mty Mu

CFTC-2.5-0.5 CFTC-2.5-0.5s CFTC-2.5-0.5r CFTC-1.8-0.5 CFTC-1.8rs-0.5 CFTC-2.5rs-0.5 CFTC-2.5rs-0.4 CFTC-2.5cs-0.5

234 234 234 234 234 234 234 234

408 486 408 336 387 486 486 523

58 58 58 71 71 58 58 58

466 466 429 442 442 466 419 466

524 524 497 513 513 524 477 524

408 486 408 336 387 486 477 524

409 498 400 372 411 480 491 550

1.00 1.02 0.98 1.11 1.06 0.99 1.03 1.05

377

Engineering Structures 182 (2019) 363–378

X. Li et al.

Base plate [7]

N1 Ft

x d0 Concrete footing

Ft

Tension Compression zone zone

[8]

b

Mu3 0.85 f'c

[9]

x

[10]

d0 h0

[11] [12]

Fig. 14. Force diagram of inner base plate connections.

[13]

Acknowledgments

[14]

This research is financially supported by the Fundamental Research Funds for the Central Universities (Grant number: 2017XKQY051).

[15] [16]

Appendix A. Supplementary material

[17]

Supplementary data to this article can be found online at https:// doi.org/10.1016/j.engstruct.2018.12.076.

[18]

References

[19] [20]

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