Seismic Behavior of Flexural Dominated Reinforced Concrete Bridge Columns at Low Temperatures

Montejo, Luis A.; Kowalsky, Mervyn J.; Hassan, Tasnim

doi:10.1061/(ASCE)0887-381X(2009)23:1(18)

Free access

TECHNICAL PAPERS

Mar 1, 2009

Seismic Behavior of Flexural Dominated Reinforced Concrete Bridge Columns at Low Temperatures

Authors: Luis A. Montejo, Mervyn J. Kowalsky, A.M.ASCE, and Tasnim Hassan, A.M.ASCEAuthor Affiliations

Publication: Journal of Cold Regions Engineering

Volume 23, Issue 1

https://doi.org/10.1061/(ASCE)0887-381X(2009)23:1(18)

PDF

Abstract

This paper presents the results from Phase II of an experimental study on the behavior of reinforced concrete bridge columns in cold seismicly active regions. Six half-scale circular reinforced concrete columns, designed to be flexural dominated, were tested under reversed cyclic loading while subjected to temperatures ranging from

- 36 ° C

(- 33 ° F)

to

22 ° C

(72 ° F)

. Four of the units tested were reinforced concrete filled steel tube (RCFST) columns and the other two were ordinary reinforced concrete columns. Results obtained reiterated the observations made in Phase I, which is that low temperatures cause an increase in the flexural strength and initial stiffness as well as a reduction in the spread of plasticity and displacement capacity of the column. Another important observation made was that the plastic hinge length is drastically reduced in the RCFST units compromising the displacement capacity of this type of column even at room temperature conditions. Current predictive models were revised and modified to account for the low-temperature effect.

Introduction

An extensive literary review in an earlier paper (Montejo et al. 2008) identified a significant lack of information on the combined effect of freezing temperatures and seismic loads on the performance of reinforced concrete (RC) columns. The same paper also presented the results from Phase I of an experimental study that consisted of the reversed cyclic testing of four ordinary reinforced concrete (ORC) circular members subjected to temperatures ranging from

- 40 ° C

(- 40 ° F)

to

20 ° C

(68 ° F)

. ORC members refer simply to conventional reinforced concrete members. It was observed that ORC members subjected to cyclic reversals undergo a gradual increase in strength and stiffness, as well as simultaneous reduction in displacement capacity as the temperature decreases. In Phase I the RC members were lightly reinforced

(ρ_{l} = 1 %)

and tested without axial load. This paper presents the results from the testing of six new specimens that were designed to more closely represent the behavior of real bridge bent columns: longitudinal steel ratios

(ρ_{l})

were between 2 and 3% and, in addition to the cyclic reversals and low temperatures, the specimens were also subjected to a constant axial load.

The purpose of the experimental program was to investigate the seismic behavior at low temperatures of ORC columns and reinforced concrete filled steel tube (RCFST) columns. In RCFST columns the steel tube also becomes the formwork during casting of the concrete. In practice, a gap is left between the steel tube end and the beam-column joint or footing (if present); so that the steel tube is only providing shear and confinement strength to the column, and not (in a direct way) flexural or axial strength (which are provided by the concrete and the longitudinal reinforcement). Some of the advantages of RCFST are that: (1) no formwork is required; (2) the whole concrete section is very well confined which, in theory, will increase the ductility capacity of the section; and (3) since the steel tube provides shear and confinement strength, a minimum number of conventional spirals or ties is required. Past research (Aboutaha and Machado 1998, 1999) has found that RCFST columns exhibit a larger displacement capacity than ordinary reinforced concrete columns when subjected to cyclic reversals and high axial loads ratios (ALRs)

> 10 %

(condition proper of tall buildings in active seismic zones). Chai et al. (1991) and Priestley et al. (1994a,b) investigated the use of steel jackets for seismic retrofit of nonductile reinforced concrete columns. The results obtained show that the columns retrofitted with steel jackets exhibited stable lateral force–displacement hysteretic response. The pattern of inelastic deformation was changed from predominantly shear deformation for the as-built columns to predominately flexural deformation for the retrofitted columns. An increase in the elastic stiffness and a reduction in the spread of plasticity were also noticed in the retrofitted columns. A reduction in the spread of plasticity was also noticed on the ORC columns tested at freezing temperatures in Phase I (Montejo et al. 2008). This was identified as the possible cause for the reduced displacement capacity of cold specimens when compared to the room temperature specimens. Therefore, it is of special interest to investigate the combined effect of low temperatures and the high confinement and stiffness provided by the steel tube in the seismic behavior of RCFST columns.

The sketch of a bridge bent and the moment distribution in the columns due to transverse lateral load is shown in Fig. 1. Each unit was intended to represent the part of the column from the cap beam to the inflection point, as represented by the shaded area in Fig. 1. The behavior of in-ground hinges, which develop in multiple column bents with continuous pile/shaft column system, is not investigated in this research. Information related to the effect of ground freezing on the behavior of in-ground hinges can be found in Suleiman et al. (2006) and Sritharan et al. (2007).

Fig. 1. Sketch of bridge bent and representative column specimen (shaded) for testing

Column Details and Test Setup

Table 1 presents the test matrix. A total of three pairs of columns were tested, two of the pairs consisted of RCFST columns and the remaining one of ORC columns. All the columns were detailed to ensure a flexural failure. The only difference between the columns in each pair was the temperature of the specimen during the test. One of the columns was tested at room temperature

22 ° C

(72 ° F)

while the other was tested at

- 36 ° C

(- 33 ° F)

. The cooling process of the cold specimens was started

26 h

before each test, where the specimen was exposed to a constant temperature of

- 40 ° C

(- 40 ° F)

. Temperatures inside the specimens were monitored by three imbedded thermocouples placed at the core and main reinforcing bars on the base of the column. Fig. 2 shows the temperatures registered during testing of specimen ORC-89C as a function of the column tip displacement. Although the temperature in the core of the column was

\sim 3 ° C

(5 ° F)

degrees warmer than at the level of the longitudinal bars, it can be seen from this figure that temperature was constant through the entire test. Similar behavior was obtained for the other two cold specimens; temperatures reported in Table 1 are the average of the temperatures recorded by the three imbedded thermocouples.

Fig. 2. Temperatures variations during testing of ORC-89C

Table 1. Test Matrix

Unit	Avg.temp	Longitudinal steel/ratio	Transverse steel/ratio	Axial load/ratio
ORC 89A	$+ 22 ° C$ $+ 72 ° F$	8#93.1%	#3 at $60 mm$ $(2.4 in.)$ 1.2%	$220 kN$ $49 kips$ 6.2%
ORC 89C	$- 36 ° C$ $- 33 ° F$	8#93.1%	#3 at $60 mm$ $(2.4 in.)$ 1.2%	$218 kN$ $49 kips$ 4.8%
RCFST 89A	$+ 22 ° C$ $+ 72 ° F$	8#93.1%	#3 at $60 mm$ $(2.4 in.)$ $9.5 mm$ $(3 ∕ 8 in.)$ th. steel tube $(1.2 + 8.5) %$	$231 kN$ $52 kips$ 5.9%
RCFST 89C	$- 36 ° C$ $- 33 ° F$	8#93.1%	#3 at $60 mm$ $(2.4 in.)$ $9.5 mm$ $(3 ∕ 8 in.)$ th. steel tube $(1.2 + 8.5) %$	$219 kN$ $49 kips$ 3.3%
RCFST 87A	$+ 22 ° C$ $+ 72 ° F$	8#72.1%	#3 at $60 mm$ $(2.4 in.)$ $9.5 mm$ $(3 ∕ 8 in.)$ th steel tube $(1.2 + 8.5) %$	$226 kN$ $51 kips$ 5.7%
RCFST 87C	$- 36 ° C$ $- 33 ° F$	8#72.1%	#3 at $60 mm$ $(2.4 in.)$ $9.5 mm$ $(3 ∕ 8 in.)$ th steel tube $(1.2 + 8.5) %$	$23 kN$ $52 kips$ 3.5%

In order to accommodate the range of temperatures desired, the columns were tested inside an environmental chamber. Due to the space limitations for testing inside the chamber, the columns were tested in a horizontal position at half the scale of the actual bridge column/pile. For all the specimens (Fig. 3), the column diameter was

457 mm

(18 in.)

, cantilever length was

1651 mm

(65 in.)

, and transverse reinforcement was in the form of spirals spaced at

60 mm

(2.4 in.)

. In addition to the spiral the RCFST units have a

457 mm

(18 in.)

outside diameter (o.d.) API-5L X52 steel pipe which is

9.5 mm

(3 ∕ 8 in.)

thick. The thickness of the pipe was selected such that the diameter-thickness ratio

(D ∕ t \sim 48)

represents that of actual pile/column bridge design practice. Accordingly, the gap between the steel tube end and the cap beam was reduced to

25 mm

(1 in.)

as typical gaps in actual bridges are

50 mm

(2 in.)

.

The support block of the specimen rested on a steel base plate that had four squat legs (Fig. 3). The interface between the footing and the base plate was filled with hydro-stone to ensure a uniform distribution of the reaction forces at the supports (Figs. 3 and 4). The legs of the base plate rested on four steel tube sections that extended through the floor of the environmental chamber to the strong floor. Four

D

35 mm

(1

3 ∕ 8 in.

). Dywidag posttensioning bars were placed through the footing and the strong floor to anchor the specimen, each bar was posttensioned to

400 kN

(90 kips)

. Lateral load was applied using an actuator with a capacity of

500 kN

(112 kips)

. The actuator was vertically connected to a steel frame which was anchored to the strong floor; an extension was designed to transmit the force from the actuator to the specimen inside the environmental chamber (Fig. 4). The axial load was applied through two cross beams: one located behind the footing and the other on top of the column. Two

D

32 mm

(1.25 in.)

threaded rods were running parallel to the column and connecting the two cross beams. Two hydraulic jacks were used to apply the axial load through the bars while two

220 kN

(50 kpis)

load cells were used to measure and control the level of axial load being applied. Both jacks were connected in parallel to a single pump in order to distribute the pressure uniformly on both sides of the column. A constant pressure valve maintained constant axial load during testing to within

\pm 10 %

of the applied load. Axial loads reported in Table 1 are the average of the summation of the values recorded by the load cells during the test. In Table 1 the axial load ratio is also calculated by taking into account the increase in concrete compressive strength at low temperatures.

Fig. 4. Photo of specimen inside environmental chamber

Deflections at different heights of the column including the point of lateral load application were recorded using string potentiometers. Curvatures within the plastic hinge region were measured using four pairs of linear potentiometers mounted on threaded rods drilled into the column. Strains in the longitudinal reinforcement, spirals, and steel tubes were measured using electrical resistance strain gauges.

Test Procedure

Subsequent to applying the axial force, all specimens were subjected to a standard cyclic loading pattern (Fig. 5), which consisted of force-controlled single cycles with increments of 25, 50, 75, and 100% of the first yield force,

F_{y}^{'}

, followed by subsequent three-cycle sets in displacement control to predefined increments of the equivalent yield displacement

Δ_{y}

. The lateral force at first yield of the longitudinal reinforcement

F_{y}^{'}

and the lateral nominal force

F_{n}

were found from a section analysis using the computer code CUMBIA (Montejo and Kowalsky 2007). The lateral nominal force is defined as the force at which the cover concrete reaches a compression strain of 0.004. The displacement corresponding to first yield

Δ_{y}^{'}

is calculated as the average of the displacements recorded in the push and pull directions during the force control cycle at

F_{y}^{'}

. The equivalent yield displacement,

Δ_{y}

, is then obtained by extrapolation between the first yield and nominal lateral forces using Eq. (1)

Δ_{y} = Δ_{y}^{'} \frac{F_{n}}{F_{y}^{'}}

(1)

Therefore, the displacements corresponding to the prescribed ductility levels are not determined until the column reaches first yield. Target forces were calculated using room temperature material properties. The room temperaturespecimens from each pair of columns were tested first and the same target forces and displacements were applied to the companion cold specimen for comparison purposes.

Material Properties

All four RCFST columns were cast from the same batch of concrete, and the two ORC columns from another batch. Compressive strength values at room temperatures are obtained from the average of three

4 in. \times 8 in.

cylinders tested on the same day of the column (Table 2). For the cold specimens, cylinders with an imbedded thermocouple were placed inside the environmental chamber from the beginning of the column specimen cooling process and tested after completion of the column test. Nevertheless, as the cylinders were tested outside the chamber, an increase in the temperature was noticed. Compressive strength at the desired temperature was then extrapolated using the equation proposed by Browne and Bamforth (1981). This expression was previously shown to provide good results to the available database of low-temperature tests (Montejo et al. 2008)

f_{c}^{'} (T) = f_{c}^{'} (20 ° C) - T w ∕ 12 0 ° C > T > - 120 ° C

(2)

where

w = concrete

moisture content. For example, for the specimen ORC-89C

f_{c}^{'} (20 ° C) = 21.4 MPa

and

f_{c}^{'} (- 28 ° C) = 26.2 MPa

, then using Eq. (2)

w = 2 %

and the compressive strength at the average temperature of the column during the test is estimated to be

f_{c}^{'} (- 36 ° C) = 21.4 - (- 36) ∕ 12 = 27.6 MPa

. Table 2 summarizes the experimental and estimated concrete compressive strengths.

Table 2. Concrete Properties

Batch	$f_{c}^{'}$ [MPa (ksi)]	$T$ [°C (°F)]
ORCs	21.4 (3.1)	$+ 22$ $(+ 72)$
	26.6 (3.8)	$- 28$ $(- 18)$
	27.6 (4.0)^a	$- 36$ $(- 33)$
RCFSTs	26.2 (3.8)	$+ 22$ $(+ 72)$
	41.3 (6.0)	$- 30$ $(- 22)$
	44.0 (6.4)^a	$- 36$ $(- 33)$

^a

Estimated.

Yield and ultimate stresses of the reinforcing steel are presented in Table 3. In this table, room temperature values were directly obtained from tension tests. Low-temperature strengths are estimated to be

\sim 11 %

larger than those exhibited at room temperature (Montejo et al. 2008).

Table 3. Reinforcing Steel Properties

Size	Stress[MPa (ksi)]
	Measured $(+ 22 ° C ∕ + 72 ° F)$		Calculated $(- 38 ° C ∕ - 37 ° F)$
	$f_{y}$	$f_{s u}$	$f_{y}$	$f_{s u}$
#9	558 (81)	703 (102)	627 (91)	778 (113)
#7	442 (64)	675 (98)	490 (71)	741 (108)
Spiral	469 (68)	655 (95)	524 (76)	723 (105)

Ordinary Reinforced Concrete Columns

Lateral Force–Displacement Response

Measured lateral force–displacement hysteresis loops for ORC-89A and ORC-89C are shown in Figs. 6(a, b), respectively. It is seen that the hysteresis loops of both units are stable up to the last cycle of

μ 8

where the topmost bar of both columns buckled. It is also noticed that the loops are not symmetric about the

x

axis; it takes more lateral force to reach a target displacement in the pull direction than in the push direction. This difference may be attributed to three facts: (1) the steel cages were slightly off center; (2) the effect of the self-weight of the column and steel fixture; and (3) the unsymmetrical distribution of stiffness in the footing due to the large metallic plate on which it rests. In order to compare the response of both specimens, an average of the measured peak forces in the first cycle at each level of ductility is used to represent the response of the column. Fig. 7 shows the average force–displacement response calculated for ORC columns. From this figure can be noticed that the effect of freezing temperatures was to increase the strength and initial stiffness of the column. If the elastic stiffness is defined at the load level required for first yield of the room temperature specimen, then the elastic stiffness of the specimen tested at

- 36 ° C

is 27% larger than the elastic stiffness of the specimen tested at

22 ° C

. The largest lateral forces reached were

326.4 kN

(at the first cycle of

μ 4

) for the cold specimen and

285.7 kN

(at the first cycle of

μ 6

) for the room temperature unit, i.e., the cold specimen exhibited 14% larger flexural strength than that of the ambient temperature. Now, if the increase in flexural strength is calculated from the average of the differences between the lateral force required to reach the given target displacements, the increase in flexural strength in the cold specimen was 16%. Displacement capacity was not affected by the low temperature since both specimens failed by buckling of the topmost bar at the same cycle and level of displacement demand. Nonetheless, failure of the cold specimen was more brittle in nature as it involved rupture of the spiral at the level that the bar buckled. Strength degradation started earlier in the cold specimen, at

μ 4

, while in the ambient temperature test it started at

μ 6

.

Fig. 6. Hysteresis loops: (✫) spiral rupture; (▲) rebar rupture; (●) rebar buckling

Fig. 7. Average first-cycle envelopes of ORC-89A and ORC-89C

Hysteretic Damping

Energy absorption capacity is evaluated by means of the hysteretic damping calculated using (Jacobsen 1930)

ξ = \frac{2}{π} \frac{A 1}{A 2}

(3)

where

A 1 = area

inside each loop; and

A 2 = area

of a rigid, perfectly plastic member with the same maximum strength and the same maximum displacement in each direction as the actual member. Note that the area-based equivalent viscous damping (AB-EVD) values calculated with Eq. (3) are not the values to use for displacement based design as they may largely overestimate the effective equivalent viscous damping for systems with high energy absorption (Chopra and Goel 2001). Appropriate levels of EVD have been calibrated for different hysteretic rules to give the same peak displacements as the hysteretic response using inelastic time history analyses (ITHA) (Dwairi and Kowalsky 2007; Grant et al. 2005). Correction factors to be applied to AB-EVD are displayed in Fig. 8 (Priestley et al. 2007); trend lines in this figure correspond to Eq. (4). Fig. 9 presents area-based damping values obtained for both specimens along with the corrected values for design. It is seen that the cold specimen exhibits slightly larger AB damping than the room temperature unit. However, when these values are corrected the difference becomes negligible. Note that the values obtained are in agreement with the equation proposed by Dwairi and Kowalsky (2007) for bridge columns [Eq. (5)]

(ITHA ∕ AB) EVD ratio = (0.53 μ + 0.8) ξ^{- (μ ∕ 40 + 0.4)}

(4)

ξ_{eq} = 50 (\frac{μ - 1}{π μ})

(5)

Fig. 8. Correction factors (Priestley et al. 2007) to be applied to AB-EVD

Fig. 9. Hysteretic damping of ORC-89A and ORC-89C

Curvature Distribution

Curvature values are calculated using the relative displacements recorded by the linear potentiometers placed on the extreme tension/compression faces of the column. The gauge length for the linear potentiometers in the base of the column includes a component due to strain penetration as it can be argued that the rotation evaluated in this point is distributed into the footing. The magnitude of this addition is obtained by optimization of the match between the experimental and theoretical moments associated with a given experimental curvature. Fig. 10 shows the match obtained for ORC-89C. The calculated curvatures can be used along with the measured tip displacement to define an equivalent plastic hinge length. Assuming that the deflection of the column after yield is attained by the formation of a plastic hinge of length,

L_{p}

, within which curvatures are equal to the base curvatures, the tip displacement

Δ

can be expressed as

Δ = Δ_{y} + ϕ_{p} L_{p} L

(6)

where

L_{p} = equivalent

plastic hinge length;

ϕ_{p} = plastic

curvature; and

L = length

from the face of the footing to the location of the applied load. The equivalent plastic hinge length is then obtained using Eq. (7)

L_{p} \sim \frac{Δ_{p}}{{\dot{ϕ}}_{p} L}

(7)

where

Δ_{p} = plastic

displacement at a given displacement ductility. The value for

Δ_{p}

was determined by subtracting the equivalent yield displacement from the displacement at the given displacement ductility, while

ϕ_{p}

was calculated by subtracting the equivalent yield curvature from the curvature at the base of the column at the given ductility. Fig. 11 compares the curvature distributions for both units. The theoretical curvature distribution for first yield of the longitudinal steel at the base of the column

ϕ_{y}^{'}

is also shown in this figure. It can be noticed that plastic curvatures at the base on the column of the cold specimen are larger when compared with the room temperatures profiles. This implies that plastic curvatures in the room temperature unit should be distributed over a larger length in order to reach a specified displacement when compared with the cold unit. This is corroborated in Fig. 12, which presents the equivalent plastic hinge lengths obtained from the experimental results. Plastic hinge lengths are calculated independently for the pull and push directions. Fig. 12 show the averages of the values obtained from both directions. Note that values of

L_{p}

obtained for the cold unit

(264 mm)

are

\sim 64 %

of the values obtained for the room temperature conditions

(412 mm)

. This reduction in the plastic hinge is strengthened by a visual inspection of the specimens after the test, where a reduction in the spread of plasticity is evident (Fig. 13). Furthermore, past research has shown that bond strength increases with cold temperatures. Shih et al. (1988) reported an increase of 76% in bond strength of bars subjected to cyclic reversals when the temperature is reduced from

20 to - 40 ° C

. This large increase in bond strength is expected to reduce the strain penetration length, thereby reducing the effective plastic hinge length,

L_{p}

, and corresponding member ductility.

Fig. 10. Moment curvature relation for ORC-89C

Fig. 11. Curvature profiles of ORC-89A and ORC-89C

Fig. 12. Equivalent plastic hinge lengths for ORC-89A and ORC-89C