Seismic behaviors of precast assembled bridge columns connected with prestressed threaded steel bar: Experimental test and hysteretic model

Abstract

Three 1/10-scale bridge pier specimens were tested under quasi-static test. The specimens included two precast specimens (PC1 and PC2) and one cast-in-place reference specimen. The two precast bridge pier specimens were connected with prestressing threaded steel bar and steel flange at the connection between precast pier column and the foundation, and non-socket assembly scheme and socket assembly scheme are adopted, respectively. They were tested to verify the seismic performance of prefabricated piers connected by prestressed threaded steel bars and steel flanges and study which assembly scheme is better for non-socketed and socketed piers. The results show that the prefabricated pier with the combination of the prestressed threaded steel bars and steel flange has higher cracking load and smaller residual displacement, which indicates that it has good service performance and good self-resetting ability. Compared with the non-socket assembly scheme, the socket assembly scheme is superior due to its higher ductility, higher overall initial stiffness, and higher energy dissipation capacity. Therefore, the prefabricated assembled pier with the socket connection scheme of the combination of the prestressed threaded steel bars and steel flange has good service performance and seismic performance. After that, a hysteretic model for the precast assembled columns was proposed, which has a good agreement with the test results.

Keywords

hysteretic model precast assembled bridge columns prestressed threaded steel bars quasi-static test seismic behaviors

Introduction

The conventional production mode of pier plays an important role in bridge construction, but the conventional cast-in-place (CIP) pier has the disadvantages of serious waste of materials, great impact on traffic environment, and difficult quality control. If the construction is in some areas with heavy traffic, it is necessary to shorten the construction time and restore traffic as soon as possible to quickly reduce the interference of bridge construction on vehicle traffic. Therefore, the use of prefabricated assemble piers has increased in recent years owing to the fast construction speed and slight traffic block (Billington et al., 2001; Zhou et al., 2017). Compared with conventional CIP piers, prefabricated piers have many inherent advantages, including not only higher construction speed, but also lower environmental impact and higher structural security (Bu et al., 2016).

Although the quality control of prefabricated pier column is superior than that of CIP concrete column on the construction site, the connection of prefabricated components seriously affects the overall performance of prefabricated assembled piers. Therefore, the connection between prefabricated components is the main factor to ensure the service performance, safety, and seismic performance (Shim et al., 2017). Therefore, reliable connection methods or technologies are necessary to improve the seismic performance of prefabricated piers (Kim et al., 2015). At present, many researchers have proposed the connection methods and technologies of pier prefabricated components and carried out experimental research and analysis. Ou (2007) regarded grouting corrugated pipe as a connection method of prefabricated columns and the footing and carried out experimental research and showed satisfactory test results. Ichinose et al. (2004), Brunesi et al. (2015), and Murcia-Delso et al. (2013) used grouting splice sleeve couplers (GSSCs) to connect prefabricated component. Compared with those of the corresponding CIP specimen, prefabricated components had considerable strength. Wang et al. (2018) analyzed the seismic performance of prefabricated piers connected by grouting sleeves and prestressing tendons through quasi-static tests. The results show that grouting sleeves have local reinforcement effect on the conventional plastic hinge area, and the prestressing tendons showed better self-resetting ability in quasi-static tests. Haber et al. (2014) and Tazarv and Saiidi (2016) carried out quasi-static tests on prefabricated piers connected by grouting sleeves alone. The test results display that the horizontal bearing capacity, ductility, and energy dissipation of prefabricated piers are very close to those of CIP piers, but the failure modes are different. Huang Yi et al. adopted steel sleeve and metal corrugated pipe as the connection of single segment assembled pier. Through the quasi-static comparative test method, the horizontal bearing capacity of assembled pier is similar to those of CIP pier, but the energy consumption and ductility of the assembled pier are worse than that of CIP pier.

Many prefabricated pier connections have been suggested by several researchers. Judging from the observed failure mode, the availability of the prefabricated pier system is mainly concerned with the connection between the foundation and the first pier section (Shim et al., 2008). In addition, segmental piers are not as good as integral piers in shear resistance at the joint of segments (Kim et al., 2010). Also, the wet joint monolithic segment is generally used to connect the high pier column with the footing in built sea-crossing bridges (Li, 2006; Wang et al., 2008). Therefore, it is still necessary to improve the integrity and seismic performance of prefabricated piers from the analysis of the weak links in the connection of prefabricated piers (Chou and Chen, 2006). New prefabricated pier connection technology is still needed to meet the needs of different construction conditions (Ge et al., 2017).

In this article, a new precast bridge pier connection method for relatively small bridge piers, using prestressed threaded steel bar and steel flange, is proposed. The steel flange is fixed at the bottom of the precast pier column by connecting the internal horizontal steel bar with the longitudinal steel bar of the precast pier column. The prestressed threaded steel bar is embedded in the reinforced concrete (RC) foundation, so the prefabricated pier connected by the prestressed threaded steel bar and the steel flange has a certain self-resetting ability. Compared with the common grouting sleeve and metal corrugated pipe grouting connection mode at present, the combined connection of prestressed threaded steel bar and steel flange is simpler in construction and easier in quality control. In this article, three 1:10-scale pier specimens were tested by quasi-static test to investigate the seismic performance of two pre-assembly schemes of non-socketed and socketed, compared with the conventional cast-in-situ pier specimens. It is expected that these two pre-assembly schemes can provide reference for the practical application of the assembled pier project. After that, a hysteretic model for three bridge pier specimens was presented and validated by tests.

Experimental program

Description of test specimens

Three bridge pier specimens, including two precast specimens (PC1 and PC2) and one CIP specimens, were constructed, and quasi-static tests were carried out. The pier specimens consist of an RC foundation of 1400 mm × 1000 mm × 500 mm with or without a 125-mm-high socket, and an RC column of 400 mm × 400 mm square section with 80-mm rounded arc. The height of three columns is 2400, 2400, and 2525 mm, respectively. The overall height of the three pier specimens is 2900 mm. Table 1 presents the properties of each specimen and Figure 1 shows the detailed dimensions of the test specimens.

Table 1.

Test specimens.

Specimen	Configuration	Column height (mm)	Overall height (mm)
CIP	Cast-in-place, connected with steel bars	2400	2900
PC1	Precast, connected with prestressed threaded steel bars, no socket column	2400	2900
PC2	Precast, connected with prestressed threaded steel bars, with socket column	2520	2900

CIP: cast-in-place.

Figure 1.

Test specimens (unit: mm): (a) specimen CIP, (b) specimen PC1, and (c) specimen PC2.

For the two precast pier specimens, PC1 and PC2, eight prestressed threaded steel bars of 18 mm diameter were embedded in the precast foundation to connect the precast column. A steel flange was embedded in the bottom of the precast column for connection. First, the precast column was assembled on the foundation by inserting the prestressed threaded steel bars into the steel flange. Eight through-hole load transducers for the steel bar were then installed on the top plate to measure the prestress. After that, shrinkage-free self-compacting micro-expansive concrete was poured into the socket of specimen PC2. Finally, the prestressed threaded steel bars were tightened to 32 kN by screw caps.

Material properties

According to the standard for test method of concrete structures GB/T 50081-2002 (2002), three 150-mm concrete cube specimens were fabricated under the same conditions with precast piers, and the quasi-static test was carried out. The average compressive strength of the concrete cube specimens was 41.33 MPa. Tensile properties of the steel bars and steel plates employed in this experimental study were tested in accordance with the ASTM A370-12 code (2012). Tensile strength of steel bars and steel plates was obtained using direct tensile tests. Table 2 presents the detailed mechanical properties of steel bars and steel plates.

Table 2.

Mechanical properties of steel bars and steel plates.

Type	Yield strength (MPa)	Tensile strength (MPa)	Ultimate strength (MPa)
Steel bars	400	438.46	653.27
Steel plates	345	261.96	406.43

Test setup and loading protocol

The quasi-static test was carried out in the Key Laboratory of Structure Engineering & Earthquake Resistance, Ministry Education, as shown in Figure 2. The loading system included one vertical hydraulic jack and one lateral actuator. The vertical hydraulic jack was fixed on the steel beam with a loading capacity of 2000 kN, which exerts a constant axial force of 310 kN on the top of the specimen to simulate the weight of the superstructure in the process of testing. The corresponding axial compressive ratio was N/(f_c·A) = 0.1, where N is the total axial force, f_c is the concrete compressive strength, and A is the column concrete section area. The lateral actuator, whose loading capacity is 1000 kN, was fixed on the reaction wall to provide reciprocating horizontal action. The location of the application force was 2220 mm up from the top of the RC foundation.

Figure 2.

Test setup and measurements (unit: mm).

The quasi-static test was carried out using hybrid control of load and displacement (Ou et al., 2009). The horizontal load of load-controlled loading is loaded at intervals of 10 kN starting from 10 kN. Each loading level is loaded for two cycles until the horizontal displacement reaches 0.075% (1.7 mm) or the pier is yielded (both occur first) and then converted to displacement-controlled loading. The displacement-controlled cyclic loading test was conducted on the basis of applied drift levels according to Specification of testing methods for earthquake resistant building (JGJ 101-1996, 1996 1997), which were initially at 0.075%, and were subsequently increased to 0.125%, 0.25%, 0.375%, 0.5%, 1%, 1.5%, 2%, 3%, 4%, 5%, 6%, 7%, 8%, 9%, 10%, and so on, and were finalized at failure. Here, the drift levels were defined as the loading displacement divided by the distance from the loading point to the top of the RC foundation. The prescribed displacements were applied on the column in three cycles for each drift ratio, as shown in Figure 3.

Figure 3.

Lateral loading histories.

Local response of three bridge pier specimens was observed by the strain measurement of the concrete within the plastic hinge location and the steel flange. In addition, displacement transducers were laid to measure lateral displacement of integral pier, joint opening, joint slipping, and the movement of the foundation by lateral loading, as shown in Figure 2.

Experimental results

Description of test failure state

The three pier specimens all showed the flexural failure mode in the plastic hinge region, as shown in Figure 4. There was no fracture or buckling of the prestressed threaded steel bar and steel plate. Three bridge pier specimens initial cracking occurred at displacement level of 5.6, 11.1, and 5.6 mm, respectively. From the comparison of the performance level of each specimen with the initial crack, the cracking load of the specimens can be increased using the connection of the prestressed threaded steel bar and steel flange. The cracking load was increased by 21% for specimen PC1 and 25.7% for specimen PC2. The cracking load of specimen PC2 is slightly higher than that of specimen PC1, but the increase is not significant. These results showed that the use of this connection method improves the performance of the precast pier specimens in the operation stage.

Figure 4.

Failure of plastic hinge zone: (a) pier failure, (b) specimen CIP, (c) specimen PC1, and (d) specimen PC2.

The final failure state of specimen CIP is observed after tests, as shown in Figure 4(b). The cover concrete in the plastic hinge zone spalled severely, yielding longitudinal reinforcement, and the stirrups broke down. With the gradual increase in loading displacement grade in the test process, existing cracks increased their width, new flexural cracks formed up to a height of about 1250 mm from the bottom of the column. Its plastic hinge region mainly concentrates at the bottom of the column, and the height of the damage area is 20 cm approximately.

The failure state of specimen PC1 after the test is shown in Figure 4(c). The height of the plastic hinge region is also about 20 cm, but the damage area is located on the upper part of the wrapped steel plate, and the wrapped steel plate has not yielded. The failure state of prefabricated bridge piers is delayed compared with conventional CIP pier specimen. It is shown that the outer steel plate at the bottom of pier column can enhance the conventional plastic hinge area. In other words, as the displacement level increases, the plastic hinge position of the specimen PC1 moves up, and severe cover concrete crushing and spalling occurred in the RC area above the outer steel plate.

The failure state of specimen PC2 at the end of the test is shown in Figure 4(d). Similarly, as to the outer steel plate at the bottom of the pier column enhances the conventional plastic hinge area, the plastic hinge moves upward as the displacement load increases. Compared with the performance of specimen PC1, specimen PC2 has larger cracking load with smaller cracking displacement. This is mainly because the shrinkage-free self-compacting micro-expansive concrete in the slot has restraint effect on the precast pier column.

In summary, the existence of steel flange increases the stiffness of conventional plastic hinge zone, and the cracking failure of piers and columns was delayed. The failure zones of the all specimens are close to each other, indicating that the outer steel plate at the bottom of the pier column moves the position of the conventional plastic hinge up to the height of the outer steel plate hoop. In addition, specimen PC2 have stronger delayed cracking load than specimen PC1.

Horizontal load–displacement hysteresis curve

The hysteretic curves obtained from the quasi-static test are shown in Figure 5. All the hysteretic loops are similar in shape and are spindle, which indicates that they have good energy dissipation capability. Compared with the specimen CIP, the specimens PC1 and PC2 have a more full hysteresis loop, indicating that the prefabricated assembled piers connected by prestressed threaded steel bars and steel flanges have better energy dissipation performance than the conventional CIP piers. In addition, these figures show that specimens CIP and PC2 exhibited stable ductility than the specimen PC1 before the displacement level of 88.8 mm. The test of three pier specimens was terminated at the displacement level of 88.8, 133.2, and 111 mm, respectively, after the loss in strength exceeded 15%. In addition, the ultimate strength of specimens PC1 and PC2 was higher than specimen CIP, but residual displacement is lower before 88.8 mm displacement level.

Figure 5.

Load–displacement hysteresis curve: (a) specimen CIP, (b) specimen PC1, and (c) specimen PC2.

Envelop curves

The comparison of envelop curves extracted from hysteretic curves of each specimen is shown in Figure 6. Characteristic points of envelop curves are shown in Table 3. All envelop curves have constant stiffness values at the initial stage of loading. It can be seen from these curves that combined connection of prestressed threaded steel bar and steel flange can contribute to lateral force. As shown in Table 3, the yield strength of specimens PC1 and PC2 increased significantly. To be specific, combined connection of prestressed threaded steel bar and steel flange can be 22.3% and 11.4% higher than specimen CIP (CIP reference column).

Figure 6.

Envelop curves of specimen.

Table 3.

Characteristic points of envelop curves.

Specimen	Loading direction	Yield load point		Maximum load point		Ultimate load point		Displacement ductility
		P_y (kN)	Δ_y (mm)	P_m (kN)	Δ_m (mm)	P_u (kN)	Δ_u (mm)
CIP	Positive	98.9	16.9	119.5	33.3	101.6	88.8	5.3
	Negative	100.3	19.7	115.3	33.3	94.6	88.8	4.2
PC1	Positive	121.1	38.4	138.4	66.6	91.6	115.1	3.1
	Negative	117.3	28.6	136.1	66.6	92.3	113.1	3.9
PC2	Positive	110.2	13.9	135.1	54.1	92.1	96.7	7.1
	Negative	119.9	19.7	140.1	66.6	74.8	85.2	4.8

CIP: cast-in-place.

Table 3 summarizes the maximum loads and their corresponding displacements. Specimens PC1 and PC2 showed a significant increase in the maximum strength. The specimens PC1 and PC2 showed similar maximum flexural strength but slightly lower than specimen CIP. The specimen PC1 showed the greatest displacement at the ultimate strength, as specimen PC1 has joint cracks at the connection between RC foundation and pier column. As a result, large horizontal displacement occurs at the top of pier column. However, the ultimate displacement of specimen PC2 is similar to that of specimen CIP due to the effect of post-poured shrinkage-free self-compacting micro-expansive concrete, which hinders the opening of the joints. Therefore, it can be considered that the socket connection scheme of specimen PC2 is equivalent to the fixed connection.

The slope of the tangent to the load–displacement curve at the origin is called the initial stiffness, and the initial stiffness is primarily for the component in the elastic phase. The initial stiffness of the three pier specimens is 21.56, 15.23, and 20.19, respectively. Compared with specimen CIP, the initial stiffness decreased by 29.3% for specimen PC1, while the initial stiffness of the specimen PC2 is almost the same as that of the specimen CIP. This is mainly because the specimen PC1 will open at the joint between the prefabricated column and the RC foundation, which weakens the lateral stiffness of the specimen PC1. However, the shrinkage-free self-compacting micro-expansive concrete in the slot of the specimen PC2 can effectively prevent the opening of the joint. It means that the prefabricated bridge pier socket connection scheme is closer to the conventional CIP pier in the elastic phase than the non-socket connection scheme.

Displacement ductility (μ = Δ_u/Δ_y) of the three pier specimens is shown in Table 3, which can be obtained from the envelop curves shown in Figure 6. The yield displacement Δ_y is determined by an equivalent bilinear curve idealized from the load–displacement relationship, and Δ_u is the displacement where the strength decreases 80% of the ultimate value (Ou et al., 2007). It can be seen that specimen PC1 exhibits weaker ductility (about 27%) than specimen CIP. However, specimen PC2 is higher (about 24%) than specimen CIP. The reason may be that specimen PC1 has joint cracks at the connection between RC foundation and pier column; it makes specimen PC1 have larger yield displacement, so the ductility of specimen PC1 is smaller. And, the ductility of the specimen PC2 is larger than that obtained from the prefabricated bridge piers connected by steel sleeve and metal corrugated pipe (Ge et al., 2017). The results show that the prefabricated pier socket prefabrication scheme with the combination of the prestressed threaded steel bar and steel flange has the best ductility.

Stiffness degradation

From the load–displacement curve, it can be seen that the secant stiffness is always changing during the loading process. With the cracking and spalling of concrete, the concrete in plastic hinge area gradually withdraws from work, and its stiffness is gradually deteriorating. The formulas for calculating secant stiffness are as follows JGJ 101-1996 (1996)

K_{i} = \frac{| + F_{i} | + | - F_{i} |}{| + X_{i} | + | - X_{i} |}

(1)

where F_i represents the load value of the first positive and negative peak points, X_i represents the displacement value of the first positive and negative peak points.

The stiffness degradation curve is shown in Figure 7. It can be seen from the figure that the stiffness degrades rapidly when displacement level is less than 33.3 mm but slows down afterward. Each specimen showed obvious difference in stiffness which indicated that different ways of connecting will have different effects. Before the displacement level of 44.4 mm, the secant stiffness value of the specimen CIP is larger than the specimen PC1 and smaller than the specimen PC2. This is mainly because the post-poured shrinkage-free self-compacting micro-expansive concrete in the slot hinders the displacement and deformation of specimen PC2. It means that the prefabricated assembled pier socket connection scheme with prestressed threaded steel bar and steel flange combination is more rigid than the conventional CIP piers in the elastic stage, which has good overall performance. Therefore, the prefabricated pier socket prefabricated assembled scheme with the combination of the prestressed threaded steel bar and steel flange has more advantages than the non-socket prefabricated assembled scheme.

Figure 7.

Stiffness degradation curves.

Energy dissipation

In order to evaluate the energy dissipation capacity, the cumulative energy dissipation is obtained by summing the area of hysteresis loop at each displacement level (Guo et al., 2016; Kwan and Billington, 2003). Note that only the first hysteresis loop at each displacement level was incorporated when computing the energy dissipation capacity. Comparative curves of the cumulative energy absorption capacity of the three piers specimens are shown in Figure 8. It can be seen from the comparison curve of energy dissipation capability, the differences between specimen PC1 and the CIP reference specimen are minor before the displacement level of 88.8 mm, while the specimen PC2 is larger than the CIP reference specimen. When the loading displacement reached the displacement level of 88.8 mm, compared with specimen CIP, the energy dissipation increased by 10.8% for specimen PC2 and decreased by 11.6% for specimen PC1. Specimen PC2 exhibits the higher energy dissipation capacity, because specimen PC2 retains better ductility. The failure of the post-poured shrinkage-free self-compacting micro-expansive concrete in the slot can dissipate part of the energy. At this time, specimens CIP and PC2 have been destroyed, but specimen PC1 still has strong energy dissipation capacity when the loading displacement reached the displacement level of 133.2 mm.

Figure 8.

Energy dissipation capability.

Residual displacement analysis

Quasi-static residual displacement Δ_res is defined as the lateral displacement of the pier when the lateral displacement reaches its maximum and the unloading is zero, as shown in Figure 9. The pier with small residual displacement is conducive to post-earthquake repair and can improve the applicability of post-earthquake piers (Pettinga et al., 2007). Therefore, the residual displacement is an important factor in the seismic performance of piers.

Figure 9.

Definition of residual displacement.

The variation of residual displacement of each specimen with loading displacement grade is shown in Figure 10. Residual displacement is chosen as the average of three cycles of residual displacement at each displacement level. By comparing the reference specimen CIP with the specimens PC1 and PC2 utilizing prestressed threaded bars, the residual displacement for specimens PC1 and PC2 was smaller than that of reference specimen CIP. When the displacement grade reaches 88.8 mm, the maximum residual displacement of specimen CIP is 63.92 mm, while the residual displacement of precast piers is less than that of specimen CIP due to the prestressed threaded bars. The residual displacements of specimens PC1 and PC2 are 83.8% and 91%, respectively, indicating that the prefabricated assembled piers connected with prestressed threaded steel bar and steel flange have better self-resetting ability than CIP piers, which is similar to the conclusion given by Wang et al. (2018) for precast bridge columns utilizing prestressing strands. On the other hand, the residual displacement for the specimen PC1 was lower than that of specimen PC2 due to the post-poured shrinkage-free self-compacting micro-expansive concrete of specimen PC2 which hinders the action of prestressed threaded steel.

Figure 10.

Residual displacement curve.

Joint sliding and joint opening

Joint sliding value at each displacement level of three pier specimens is illustrated in Figure 11. It could be found that the sliding of prefabricated piers at the connection between pier columns and the RC foundation was larger than that of conventional CIP piers. Compared with specimen PC1, the specimen PC2 has smaller sliding. The main reason is that the post-poured shrinkage-free self-compacting micro-expansive concrete hinders the joint sliding of specimen PC2 with socket prefabricated scheme.

Figure 11.

Joint sliding.

Joint opening value at each displacement level of the three specimens is illustrated in Figure 12. It could be found that the joint opening of specimen PC2 at the connection between pier columns and the RC foundation was slightly larger than that of conventional CIP piers. However, the joint opening of specimen PC1 at the connection between pier columns and the RC foundation was very large. The joint opening at the connection between pier columns and the RC foundation may be weakened in a certain extent with post-poured shrinkage-free self-compacting micro-expansive concrete in socket. In summary, specimen PC2 with socket precast assembly scheme is closer to the fixed state of conventional cast-in-situ piers at the connection between pier columns and the RC foundation.

Figure 12.

Joint opening.

Analysis of lateral displacement of pier column

From the displacement transducers deployment vertically along the pier column, the variation of lateral displacement of pier column along height direction was clearly observed as shown in Figure 13. When the displacement level is small, the deformation of pier columns seems regular. With the increase in displacement level, the damage of pier columns is serious increasingly. In the plastic hinge region, the lateral displacement of pier columns changes irregularly.

Figure 13.

The lateral displacement variation curve along the height of pier column: (a) specimen CIP, (b) specimen PC1, and (c) specimen PC2.

For specimen CIP, as shown in Figure 13(a), the lateral displacement variation curve along the height of pier column presents irregular shape at the bottom of the pier. It shows that the specimen CIP destroys most severely at the bottom of the pier column, thus forming a plastic hinge. Nevertheless, for the specimens PC1 and PC2, as shown in Figure 13(b) and (c), the irregular area of the lateral displacement curve along the height of the pier is higher than that of the specimen CIP. It is explained that the steel flange at the bottom of the pier column strengthens the bottom area of the pier column and makes the original plastic hinge area move up. Compared with the lateral displacement variation curve of specimen PC1, specimen PC2 appears more regular in the bottom area of pier column. It shows that the socket precast assembling scheme can effectively hinder the rotation of the steel flange area at the bottom of the pier column and make the precast assembling pier closer to the fixed connection at the connection between the pier column and the foundation.

Hysteretic model

Summing up the experimental results and research results of other scholars (Wang et al., 2010), an improved hysteretic model which is applicable for prefabricated assembled piers was proposed, as shown in Figure 14. The model can accurately represent the self-resetting ability, nonlinear behavior, stiffness degradation based on a set of rules which depend on the properties of the piers, and the history of loading. Figure 15 displays the idealized trilinear model. The first branch of the trilinear model is elastic and the stiffness is calculated as k₁ = P_y/Δ_y. The second branch represents the stiffness degradation caused by the plastic hinge concrete cracking and steel yielding, and the plastic stiffness is calculated as k₂ = (P_m − P_y)/(Δ_m − Δ_y). The third branch represents negative stiffness with the development of the displacement and negative stiffness k₃ has a certain correlation with the initial stiffness k₁ of the elastic phase of the specimen.

Figure 14.

Modified hysteresis model.

Figure 15.

An idealized trilinear model.

The unloading stiffness gradually degenerates with the increase in displacement. In order to estimate the unloading stiffness k₄, the calculated values were expressed in terms of the elastic stiffness (k₁) and the displacements of the yield point and peak point (Δ_p, Δ_y). The correlation between them can be calculated by equation (2)

k_{4} = a {(\frac{Δ}{Δ_{y}})}^{b} k_{1}

(2)

where the value of coefficient a = 1.20, 1.22, 1.44 and b = −0.303, −0.205, −0.438 for specimen CIP, specimen PC1, specimen PC2 had been proposed on the basis of the experiment results of the tested specimens. The comparison of hysteretic model and experimental data of three specimens is shown in Table 4.

Table 4.

Comparison of hysteretic model and test data.

Specimens	k₁ (kN/mm)		k₂ (kN/mm)		k₃ (kN/mm)		k₄ (kN/mm)
	Test	Model	Test	Model	Test	Model	Test	Model
CIP	5.27	5.47	1.19	1.18	−0.36	−0.35	−	5.47
PC1	3.47	3.62	0.59	0.55	−0.63	−0.63	−	3.62
PC2	6.85	6.99	0.51	0.52	−0.31	−0.31	−	6.99
Specimens	P_y (kN)		Δ_y (kN)		P_m (kN)		Δ_m (kN/mm)
	Test	Model	Test	Model	Test	Model	Test	Model
CIP	99.61	96.49	18.3	18.9	117.4	111.43	33.3	33.31
PC1	116.21	112.09	33.5	33.4	137.25	130.24	66.6	66.63
PC2	115.05	114.93	16.8	16.3	137.6	131.29	60.35	60.38

CIP: cast-in-place.

The hysteretic loop rules of the specimens are described as follows: (1) When the pier specimen is in the yield state, the loading and unloading are all along the elastic stage of the skeleton curve. The loading in positive and negative directions moves along lines 0-1 and 0-2, respectively. And unloading returns along the original path. (2) When the pier specimen exceeds the yield point (point 1) but has not reached the peak point (point 7), the positive directions loading along the skeleton curve to point 3, then unloaded to point 4, and the unloading stiffness is determined according to formula (2). When the negative directions loading reaches the yield point 2, it travels along the skeleton curve to the point 5 unloading. The unloading stiffness is also determined according to formula (2). After the reverse loading is unloaded to zero point, the positive loading points to the maximum displacement point of the last positive loading. (3) When the load reaches the peak value, the path of the loading and unloading are the same as (2), that is, 0-1-8-9 and 9-5-11-12 in positive and negative directions, respectively.

Figure 16 shows the comparison between the experimental results and proposed hysteretic model. Good agreement is observed which indicated the validity of the proposed hysteretic model.

Figure 16.

Comparison of hysteretic curves between test and theory: (a) specimen CIP, (b) specimen PC1, and (c) specimen PC2.

Conclusion

In this article, the quasi-static test of two connection schemes of prefabricated assembled piers, of non-socket and socket, which are joined by prestressed threaded steel bar and steel flange, was carried out and compared with the conventional CIP pier specimens under the same parameters. The failure mode, energy dissipation, ductility, and residual displacement of the specimens are analyzed in detail. The main conclusions are as follows:

Prefabricated and assembled piers connected by prestressed threaded steel bar and steel flange have larger cracking load than conventional CIP piers, which delays the damage of CIP piers and improves their service performance. The steel flange at the bottom of prefabricated pier column strengthens the initial stiffness of the conventional plastic hinge zone, which forms a new plastic hinge zone on the upper part of the outer steel plate, and makes the original plastic hinge zone move up the height of a steel plate. It shows that the wrapped steel plate delays the damage of the pier.

Compared with the reference specimen CIP, the ductility of the specimen PC2 increased by 24%, while the ductility of the specimen PC1 decreased by 27%. The energy dissipation capacity of the specimen PC2 increased by 10.8% compared with the specimen CIP, and the energy dissipation capacity of the specimen PC1 has been reduced by 11.6%. In terms of residual displacement, the specimens PC2 and PC1 were reduced by 9% and 16.2%, respectively. The test results show that the prestressed threaded bars can improve the self-resetting ability, and the comprehensive consideration of the socket connection scheme is superior to the non-socket connection scheme.

A hysteretic model for prefabricated assembled piers connected by prestressed threaded steel bar and steel flange was proposed with detailed loading and unloading rules, and good agreement between the experimental results and model data was observed.

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 work was supported by the National Natural Science Foundation of China (grant no. 51878550) and the Education Department Project of Shaanxi Provincial Government (grant no. 18JS066).

ORCID iD

Qifang Xie

References

American Society for Testing and Materials (ASTM) A370-12 (2012) Standard Test Methods and Definitions for Mechanical Testing of Steel Products. West Conshohocken, PA: ASTM.

Billington

Barnes

Breen

(2001) Alternate substructure systems for standard highway bridges. Journal of Bridge Engineering 6(2): 87–94.

Brunesi

Nascimbene

Deyanova

, et al. (2015) Numerical simulation of hollow steel profiles for lightweight concrete sandwich panels. Computers and Concrete 15(6): 951–972.

Lee

, et al. (2016) Hysteretic modeling of unbonded posttensioned precast segmental bridge columns with circular section based on cyclic loading test. Journal of Bridge Engineering 21(6): 04016016.

Chou

Chen

(2006) Cyclic tests of post-tensioned precast CFT segmental bridge columns with unbonded strands. Earthquake Engineering & Structural Dynamics 35: 159–175.

GB/T 50081-2002 (2002) Standard for Test Method of Mechanical Properties on Ordinary Concrete. Beijing, China: Ministry of Construction of the People’s Republic of China (in Chinese).

Yan

Wang

(2017) Experimental study on seismic performance of prefabricated and assembled piers of rail transit connected by mechanical sleeves. Earthquake Engineering & Structural Dynamics 37(6): 143–153.

Guo

Cao

, et al. (2016) Cyclic load tests on self-centering concrete pier with external dissipators and enhanced durability. Journal of Structural Engineering: ASCE 142(1): 04015088.

Haber

Saiidi

Sanders

(2014) Seismic performance of precast columns with mechanically spliced column-footing connections. ACI Structural Journal 111(3): 639–650.

10.

Ichinose

Kanayama

Inoue

, et al. (2004) Size effect on bond strength of deformed bars. Construction and Building Materials 18(7): 549–558.

11.

JGJ 101-1996 (1996) Specification of Test Method for Earthquake Resistant Building. Beijing, China: China Academy of Building Research (in Chinese).

12.

Kim

Moon

Kim

, et al. (2015) Experimental test and seismic performance of partial precast concrete segmental bridge column with cast-in-place base. Engineering Structures 100: 178–188.

13.

Kim

Lee

Kim

, et al. (2010) Performance assessment of precast concrete segmental bridge columns with a shear resistant connecting structure. Engineering Structures 32: 1292–1303.

14.

Kwan

Billington

(2003) Unbonded posttensioned concrete bridge piers. II: seismic analyses. Journal of Bridge Engineering 8(2): 102–111.

15.

(2006) Calculating Method for Design of Externally Prestressed Concrete Bridges. Shanghai, China: Tongji University.

16.

Murcia-Delso

Stavridis

Shing

(2013) Bond strength and cyclic bond deterioration of large-diameter bars. ACI Structural Journal 110(4): 659–670.

17.

(2007) Precast Segmental Post-tensioned Concrete Bridge Columns for Seismic Regions. Buffalo, NY: The State University of New York.

18.

Chiewanichakorn

Aref

, et al. (2007) Seismic performance of segmental precast unbonded posttensioned concrete bridge columns. Journal of Structural Engineering: ASCE 133(11): 1636–1647.

19.

Wang

Tsai

, et al. (2009) Large-scale experimental study of precast segmental unbonded posttensioned concrete bridge columns for seismic regions. Journal of Structural Engineering: ASCE 136(3): 255–264.

20.

Pettinga

Christopoulos

Pampanin

, et al. (2007) Effectiveness of simple approaches in mitigating residual deformations in buildings. Earthquake Engineering & Structural Dynamics 36(12): 1763–1783.

21.

Shim

Chung

Kim

(2008) Experimental evaluation of seismic performance of precast segmental bridge piers with a circular solid section. Engineering Structures 30(12): 3782–3792.

22.

Shim

Lee

Park

, et al. (2017) Experiments on prefabricated segmental bridge piers with continuous longitudinal reinforcing bars. Engineering Structures 132: 671–683.

23.

Tazarv

Saiidi

(2016) Seismic design of bridge columns incorporating mechanical bar splices in plastic hinge regions. Engineering Structures 124: 507–520.

24.

Wang

Chen

Zhao

, et al. (2010) Experimental study on SRC columns with high rate of encased steel and their restoring force model. Earthquake Engineering & Structural Dynamics 30(4): 57–65.

25.

Wang

Chang

, et al. (2008) Large-scale seismic tests of tall concrete bridge columns with precast segmental construction. Earthquake Engineering & Structural Dynamics 37: 1449–1465.

26.

Wang

, et al. (2018) Quasi-static cyclic tests of precast bridge columns with different connection details for high seismic zones. Engineering Structures 158: 13–27.

27.

Zhou

Lee

(2017) Bond-slip responses of stainless reinforcing bars in grouted ducts. Engineering Structures 141: 651–665.