Tenth U.S. National Conference on Earthquake EngineeringFrontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska 10NCEE

IMPACT OF DIAMETER TO THICKNESS RATIO ON THE SEISMIC BEHAVIOR OF

REINFORCED CONCRETE FILLED STEEL TUBES

N.K. Brown1, M.J. Kowalsky2, and J.M. Nau2

ABSTRACT Reinforced concrete filled steel tubes (RCFSTs) are commonly used as bridge pipe piles in high seismic regions. The pipe-piles consist of reinforced concrete encased in a steel tube. The steel tube is used as a permanent casing which eases construction. The concrete is confined by the steel tube, increasing the compressive strength, and the concrete core prevents the steel tube from buckling inward. Tests were performed on twelve large scale RCFSTs, seven of which are reported here focusing on the effect of varying D/t ratio. The pipe-piles were subjected to reversed cyclic four-point bending with a constant moment region centered in the pile. The large scale specimens consisted of outer diameters of 20 to 24 inches and diameter-to-thickness ratios between 33 and 192. The effect of D/t ratio on the initiation of buckling and rupture, ductility, and hysteretic damping are presented.

1 Engineer, AECOM. Formerly, MS Student, Dept. of Civil, Construction, and Environmental Engineering, North Carolina State University, Raleigh, NC 27695 2 Professor, Dept. of Civil, Construction, and Environmental Engineering, North Carolina State University, Raleigh, NC 27695 Brown, N.K, Kowalsky, M.J., Nau, J.M. Impact of Diameter to Thickness Ratio on the Seismic Behavior of Reinforced Concrete Filled Steel Tubes. Proceedings of the 10th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

Impact of Diameter to Thickness Ratio on the Seismic Behavior of

Reinforced Concrete Filled Steel Tubes

N.K. Brown1 , M.J. Kowalsky2, and J.M. Nau2

ABSTRACT Reinforced concrete filled steel tubes (RCFSTs) are commonly used as bridge pipe piles in high

seismic regions. The pipe-piles consist of reinforced concrete encased in a steel tube. The steel tube is used as a permanent casing which eases construction. The concrete is confined by the steel tube, increasing the compressive strength, and the concrete core prevents the steel tube from buckling inward. Tests were performed on twelve large scale RCFSTs, seven of which are reported here focusing on the effect of varying D/t ratio. The pipe-piles were subjected to reversed cyclic four-point bending with a constant moment region centered in the pile. The large scale specimens consisted of outer diameters of 20 to 24 inches and diameter-to-thickness ratios between 33 and 192. The effect of D/t ratio on the initiation of buckling and rupture, ductility, and hysteretic damping are presented.

Introduction Concrete filled steel tubes are used in a variety of structures including bridges, high rise buildings, and power plants. They are especially used in structures where large moments and displacements must be resisted, such as in high seismic regions, and present several possible advantages over conventional reinforced concrete or hollow steel tube construction. These advantages include simplified construction, where steel tubes can serve as permanent formwork, and improved non-linear behavior due to high levels of confinement. Concrete filled steel tubes (CFSTs) have been constructed for over 50 years and have become increasingly more common in a variety of structures. There has been substantially less research on reinforced concrete filled tubes (RCFSTs) where internal reinforcement is present. The study of RCFTSs is important for several reasons. First, there is little research available on thinner walled pipes with large diameter to thickness (D/t) ratios in which the internal reinforcement could have more of an impact on behavior, since it represents a larger proportion of the total steel area. Also, internal reinforcement in thinner walled pipes is necessary to enhance overall resistance to corrosion. In the case of bridge structures, RCFSTs are used in both drilled shaft and driven pile 1 Engineer, AECOM. Formerly, MS Student, Dept. of Civil, Construction, and Environmental Engineering, North Carolina State University, Raleigh, NC 27695 2Professor, Dept. of Civil, Construction, and Environmental Engineering, North Carolina State University, Raleigh, NC 27695 Brown, N.K, Kowalsky, M.J., Nau, J.M. Impact of Diameter to Thickness Ratio on the Seismic Behavior of Reinforced Concrete Filled Steel Tubes. Proceedings of the 10th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

applications. In some drilled shaft applications, the member below grade may be of a larger diameter than the column above grade. In other drilled shaft applications, the diameter remains constant below and above grade. In driven pile applications, the member below grade may be part of a pile supported footing, or could be continuous above grade to the bridge cap beam. While the methods of construction are different for the above cases, each application concludes with a composite system that may develop plastic hinges below grade. Above grade, RCFSTs are often terminated just short of the bridge cap beam where a connection is made using conventional reinforced concrete. In such cases, the steel tube acts only in confinement whereas in the maximum moment region below grade, the tube is subjected to both lateral and longitudinal strains. It is the goal of this research to study the non-linear seismic behavior of composite RCFSTs. The variables considered in this study are D/t ratio of the steel tube and internal reinforcement content. Of specific interest in this paper is the impact of the D/t ratio on: (1) The onset of tube wall local buckling; (2) Failure, as defined by fracture of the tube material; and (3) Hysteretic damping. These goals were accomplished through a combination of experimental and analytical methods, although the focus in this paper is on the outcome of the large scale experimentation.

Past Research

Literature on the subject of “Concrete Filled Steel Tubes” is vast and dates back to the 1950s, as noted by Park et al. [1] and Knowles and Park [2]. The focus of past research can be divided into the following broad categories: (1) Impact of confinement (including the effect of multi-axial stresses on response); (2) Issues related to composite action and strain compatibility under axial and flexural response; and (3) Member level response (including flexural strength and stiffness, shear strength, and ductility) as effected by geometric variables including D/t ratio and internal steel. A detailed literature review on all of these areas can be found elsewhere [3], [4]. Summarized in this paper is past work on the impact of D/t and internal reinforcement on the seismic behavior of both CFSTs and RCFSTs. It has been well established that slenderness parameters influence the capacity to develop full moment strength in hollow steel tube sections. For example, the current AASHTO Guide Specifications for LRFD Seismic Bridge Design [5] require a hollow pipe maximum D/t of 0.09E/Fy for essentially elastic elements, and 0.044E/Fy for ductile elements. For typical values of E and Fy (29,000 ksi and 50 ksi, respectively) this implies D/t limits of 52 (elastic) and 26 (ductile). While these limits may be suitable for hollow sections, they are likely overly-conservative for concrete-filled sections [6]. Boyd et al. [6] investigated the flexural behavior of five CFSTs subjected to a constant axial load and reversed cyclic load. All specimens were 8 inches in diameter and had D/t ratios of 73 and 107. They found that the steel shell increased energy dissipation when compared to conventional reinforced concrete columns. Interestingly, the piles with the higher D/t ratio sustained larger deformations. This is not what was expected and the researchers attributed this to differences in material ductility [6]. Bruneau and Marson [7] tested four CFST bridge columns with diameters ranging from 12.76 in to 15.98 in and D/t ratios of 31, 42, 51, and 64. Unlike the tests performed by Boyd et al. [6], the pipes with lower D/t ratios sustained higher deformations.

The thicker pipe (D/t = 34) buckled at 3% drift, one cycle later than the tests with D/t ratio of 42 and 64, which showed signs of buckling at 2%. The test with a D/t of 51 showed signs of buckling at 0.75% drift, and the researchers noted that this was unexpected but offered no insight into the inconsistency when compared to the other tests. The tests with D/t ratios of 42 and 64 ruptured at a drift of 7%, while the testing equipment stopped working during the test with a D/t ratio of 34. As a summary, the thicker pipes sustained local buckling at larger deformations, but all ruptured at approximately the same drift [7]. This outcome will be revisited later in this paper. Numerous studies have been conducted on very small scale CFSTs. Many of these tests examined the impact of D/t, loading history, and loading conditions. However, due to their size, in many cases as small as 1.7” in diameter, they are not discussed further here. The interested reader is referred to Brown et al. [4], for a detailed review of this work. Only one prior study [1] has been uncovered on the seismic behavior of RCFSTs. The primary variable in that study was D/t ratio, which was varied from 34 to 214. The specimens were 142” long with an outer diameter that varied from 10.63” to 17.7”. The piles were subjected to two cycle set reversed cyclic loading until rupture of the steel pipe or until the limitations of the equipment were reached. A point load was applied at mid-span of the pile through a concrete block which wrapped around the pipe. The mode of failure of the piles was similar for all D/t ratios where the steel pipe sustained local buckling near the loading point. The buckles increased in size until rupture of the steel pipe or maximum capacity of the equipment. The thickest pile (D/t = 34) buckled at ductility four, and the remainder of the piles buckled at a ductility of two. The height of the buckles was relatively larger in the thinner pipes (with a higher D/t ratio) than in the thicker pipes. However, the affected area or “length” of the buckles was greater in the thicker pipes than the thinner pipes [1]. All of the tests exhibited good energy dissipation, although slight pinching was noted due to concrete cracking and the closing of these cracks under reversals of loading. The “pinching” was worse in the specimens with thin walls than for those with thick walls. This difference indicated that the thin-walled piles, with a lower percentage of steel, created a response substantially similar to reinforced concrete than thicker walled piles [1]. Electrical resistance strain gages were placed on the surface of the steel tube and on the internal reinforcing cage in a rectangular rosette which allowed the researchers to obtain the lateral strains, longitudinal strains and section curvatures. A maximum longitudinal strain of about 2.5% was seen in all tests, regardless of D/t ratio.

Overview of Experimental Program Twelve 30’ long by 20” to 24” diameter tests on RCFSTs were part of this research program, although only the seven tests where D/t was varied from 33 to 192 and the internal reinforcement was constant are reported in this paper. The remaining tests had variable levels of internal reinforcing steel, which did not impact response in any appreciable manner. The interested reader is referred to Brown et al. [4] for more details on those tests. The test setup consisted of two point loads, spaced six feet apart, in the center of the specimen to create a constant moment region. The actuators which applied these loads were pushed and pulled simultaneously at the same rate. One side of the pile was supported with a “pinned” connection which allowed rotation of the pile at the support. The other side of the pile was supported with a “roller” connection with a mechanism created by connecting two steel pieces with a small steel pin

allowing rotation and horizontal displacement of the pile. A rendering of the test setup is shown in Figure 1.

Figure 1. RCFST Four point bending reversed cyclic loading test setup. Multiple instrumentation systems were utilized in the test: (1) strain gages, (2) a non-contact three-dimensional position measurement system and (3) string potentiometers. Strain gages were placed on the surface of the steel pipe in the constant moment region in the longitudinal and transverse directions. Since electrical resistance strain gages tend to be unreliable at large strains, a non-contact three dimensional position measurement was used throughout the test as well. The Optotrak Certus HD system has been proven to be very effective in monitoring large deformations in structural elements during testing. A paper by Goodnight et al. [8] in these proceedings discusses how the system has been utilized for seismic research applications at NCSU. A grid of over 200 targets was placed in the constant moment region to measure the movement of each point throughout the test. The position data collected from the targets is then used to calculate strain, curvature, rotation, and displacement. String pots were placed in the constant moment region to measure the central displacement of the specimen (providing a point of comparison with the non-contact system). A drawing showing a typical instrumentation scheme for the non-contact system is shown in Figure 2, and specific details of all test specimens are shown in Table 1.

Figure 2. Typical instrumentation scheme for non-contact LEDs.

Actuators

Steel “Pinned” Support

Steel “Roller” Support

Specimen

Test

No

Diameter

(in)

Nominal

Thickness

(in)

D/t

Ratio

Longitudinal Rebar Transverse Rebar

(Spiral)

Number Size Ratio (ρ) Size Spacing

(in)

1 24 0.5 48 12 US #7 1.60% #3 12

2 24 0.125 192 12 US #7 1.60% #3 12

3 24 0.1875 128 12 US #7 1.60% #3 12

4 24 0.375 64 12 US #7 1.60% #3 12

5 24 0.281 85 12 US #7 1.60% #3 12

6 20 0.61 33 12 US #6 1.69% #3 12

7 20 0.125 160 12 US #6 1.69% #3 12

Table 1. Test matrix.

Loading History Before yield, the loading of the pile was force-controlled and separated into four increments: (1) quarter yield force, (2) half yield force, (3) three-quarter yield force, and (4) yield force. After yield, the loading of the pile was displacement controlled, using the equivalent yield displacement, i.e. displacement ductility of 1, as the basis for additional increments, which consisted of Ductility 1, 1.5, 2, 3, 4, 5, 6, and 7. The equivalent yield displacement was obtained by extrapolating the first yield displacement (defined as the displacement when the extreme pipe fiber reached yield) to the nominal moment (defined as the moment when the extreme fiber concrete compression strain is 0.004). The specimens were tested in reversed cyclic, three-cycle set loading until rupture. The damage in each of the tests was located in the constant moment region and no damage was observed outside of this region. The mode of failure was the same for all of the tests: the steel tube suffered outward local buckling on the compression side of the pipe, which increased in size during further loading until fracture under tension. The primary difference between the tests was the shape of the buckles, and the strain at which they occurred as will be discussed in the following sections. Thin Walled Pipes (D/t >= 82) As an example of typical behavior for pipes with large D/t ratios, the response of Test 2, which had a D/t ratio of 192 is discussed in this section. The specimen showed visible signs of local buckling on the underside of the pile during the third pull cycle of ductility 1.5, at a mid-span displacement of 3.06 inches. In the reverse cycle, the first push of ductility two, local buckling also developed on the top of the pile with the most apparent buckles located under each loading point where the moment gradient changes. Smaller ripples formed throughout the constant moment region. The buckles on both sides of the pile continued to increase with increased

application of lateral deformation to the pile, although strength was sustained. The pile ruptured during the second pull of ductility six at a mid-span displacement of 12.5 inches. The progression of the buckle that ultimately ruptured is shown in Figure 3. The pile continued to be pushed and pulled after the rupture of the pipe to complete the ductility six three-cycle set. During the next pull after rupture, the pipe began to open and it was observed that the concrete under the rupture was not crushed and was still intact. A significant outcome is that the onset of local buckling did not immediately precipitate failure as the test unit was able to sustain inelastic deformations until a displacement ductility if 6 prior to fracture and loss of strength. This will be discussed further later.

Figure 3. Progression of damage for a thin wall pipe (D/t = 192), Test 2. Thick Walled Pipes (D/t <= 64) As an example of typical thick wall pipe behavior, the pipe with the smallest D/t ratio (Test 6, D/t ratio = 33) is discussed in this section. This test unit did not show any signs of damage until much higher levels of ductility and displacement. The pile showed visible signs of buckling at ductility five corresponding to a mid-span displacement of 13 inches. Buckling initiated in the first push of ductility five on the top of the pile, the compressive region, and when the cycle was reversed (the first pull) the underside of the pile also developed small signs of buckling. During the initiation of buckling, buckles only began to form under each loading point and did not form ripples throughout the constant moment region as was the case for the thin wall pipe. The buckles continued to increase in size as the mid-span displacement increased in ductility six. Although a couple secondary buckles did form during this ductility cycle, the significant buckles were still located under each loading point. During the first pull cycle of ductility seven, the buckles that had grown from the previous push cycle did not straighten and fractured as a consequence. Unlike the rupture in the thin-walled pipes, the rupture of the pipe occurred on the peak of the buckle, not in the valley as observed in thinner walled tests. The progression of buckling until rupture is shown in Figure 4, while the profile of displacement during the ductility

Ductility 2, Δ=2.04 in Ductility 3, Δ= 3.06 in Ductility 4, Δ= 4.08in Two

Ductility 5, Δ=6.12 in Ductility 6, Δ=8.17 in Rupture at Ductility 6 pull 2

seven push cycle can be seen in Figure 5. The buckles formed at a higher ductility level than in the thin-walled tests but once formed, increased in size quickly and reached the same height as those in the thin-walled tests, however due to the thickness of the pipe they were not as narrow.

Figure 4. Progression of damage for a thick wall (D/t = 33), Test 6.

Figure 5. Profile of displaced shape during push cycle of ductility 7 for Test 6.

Ductility 6 push 1, Δ = 15.6in Ductility 6 push 2, Δ = 15.6in Ductility 6 push 3, Δ = 15.6in

Ductility 7 push 1, Δ = 18.2in Ductility 7 – pull 1

Ductility 5 push 1, Δ = 13 in Ductility 5 push 2, Δ = 13 in Ductility 5 push 3, Δ = 13 in

Summary of Hysteretic Responses The force per actuator vs. midspan displacement hysteretic responses for tests 1-7 are shown in Figure 6. The following observations can be made from these graphs: (1) Energy dissipation is greater for pipes with smaller values of D/t; (2) The earlier onset of local buckling for pipes with larger D/t ratios does not correlate to an earlier onset of failure as all pipes fractured at displacement ductility levels of 6 or 7; (3) Although not shown here, the amount of internal steel did not alter the failure mode or effect response in any significant manner.

Figure 6. Sample force displacement hysteretic responses.

Design Considerations Local Buckling and Ultimate Limit States From the Optotrack data, the tensile strains prior to buckling and fracture can be identified. Tensile strains are deemed to be important to characterize local buckling as extension of the pipe under tension results in the pipe becoming the sole source for compression zone stability under reversal of loading until concrete cracks close. As a consequence, the degree to which tensile strains can be sustained is directly associated with the ability of the pipe wall to resist local buckling under compression. As was previously discussed, the onset of local buckling did not precipitate failure as the member continued to sustain lateral load until pipe fracture eventually occurred. Nonetheless, it is an important limit state as an engineer may wish to consider it as a damage control level, exceedance of which becomes difficult to repair. With regard to fracture, it is evident that it defines an ultimate limit state, beyond which the member starts to loose lateral load carrying capacity. Analysis of the Optotrack data has resulted in Figure 7 which represents these two limit state strains as a function of D/t ratio. The tension strain limit prior to local buckling under subsequent compression loading is shown as a dashed line and is essentially linear with

increasing D/t ratio. As expected, lager D/t ratios result in lower tensile strain limits prior to local buckling. The surprising result is that fracture was delayed in some cases by several ductility levels. Upon further examination, it can be postulated that such a response is a consequence of the comparatively lower strain gradient across the pipe wall thickness for larger D/t ratio. Consider two pipe thicknesses of 1/8” and ½”, with all other variables the same. The thinner wall pipe will sustain a lower strain gradient and curvature across the wall thickness once buckling occur, while the thicker wall pipe will be subjected to higher plastic strains across the thickness in the buckled regions. As a consequence, once the buckles form in the thicker wall pipe, fracture is imminent. Equation 1 represents the recommended expression for the tensile strain limit prior to the onset of local buckling, while a constant value of 2.5% can be adopted for fracture, which is consistent with the observation of Park et al. [1].

Figure 7. RCFST Tensile strain limit states as a function of D/t.

Equivalent Viscous Damping

Equivalent viscous damping is an important variable in the Direct Displacement-Based Design (DDBD) method [9]. It is also a useful measure of energy dissipation capacity. The equivalent viscous damping consists of two components: (1) hysteretic damping and (2) elastic damping. Jacobsen [10] proposed an expression for hysteretic damping based on equating the energy absorbed by hysteretic steady-state cyclic response at a given displacement level. In the expression, the area of the hysteretic loop is directly proportional to the hysteretic damping of the system. The complete force-displacement loops for all ductility levels can be extracted from Figure 6. Examination of the two extremes in Figure 6 (D/t of 32 and 192) indicate the profound effect of D/t on the enclosed area of the hysteretic response, where the thin wall pipes exhibit some pinching and proportionally smaller amounts of energy dissipation. Jacobsen’s hysteretic damping was calculated for each loop of each test. The resulting data was then corrected, and 5% tangent stiffness proportional damping added. Guidelines for this process can be found in Priestley et al. [9] and specific calculations for these tests found in King et al. [4]. The final outcome of these calculations is the equivalent viscous damping as a function of displacement ductility and D/t ratio, as shown in Figure 8 and Equation 2.

Prior to Buckling

Prior to Rupture

00.0050.01

0.0150.02

0.0250.03

0.0350.04

0 50 100 150 200Long

itud

inal

Ten

sile

Str

ains

D/t Ratio

Figure 8. Proposed equivalent viscous damping relationship for RCFST

Conclusions and Recommendations

A total of twelve large scale reversed cyclic loading tests were conducted, seven of which are reported in this paper where the primary variable was D/t ratio. The impact of D/t ratio on local buckling of the pipe wall and equivalent viscous damping was identified (Equations 1 and 2). A key observation was that D/t ratio had a profound impact on local buckling, but no impact on the ultimate limit state as defined by pipe fracture. εbuckling =0.021 - 0.00011(D/t) (1) = 0.05 + (− . ( ⁄ ) + . ) ∗ ( − )/( + . ) (2)

Acknowledgements

This research was supported by the Alaska Department of Transportation and Alaska University Transportation Center. In addition, the contributions of Elmer Marx of Alaska DOT were instrumental during the conduct of the research.

References 1. Park, R.J.T., Priestley M.J.N., Walpole, W.R. (1983). “Seismic performance of steel encased reinforced

concrete bridge piles” Bulletin of the NZ Society for Earthquake Engineering, V16#2, June, pp123-140. 2. Knowles, R.B. and Park, R. (1969). “Strength of concrete filled steel tubular columns.” ASCE Journal of

Structural Engineering, Vol. 95 #12, pp 2565-2586. 3. Hajjar, J. (2000). “Concrete-filled steel tube columns under earthquake loads” Prog. Struct. Engng Mater. 2000;

2: 72–81 4. Brown, N.R., Kowalsky, M.J., and Nau, J. (2013). “Strain Limits for Concrete Filled Steel Tubes”. Final

Report, AKDOT-AUTC. 5. AASHTO (2011). “Guide Specifications for LRFD Seismic Bridge Design”, AASHTO. 6. Boyd, F.P., Cofer, W.F, McLean, D.I. (1995). “Seismic performance of steel-encased concrete columns under

flexural loading”. ACI Structural Journal, May, pp355-364. 7. Bruneau, M., and Marson, J. (2004). “Seismic design of concrete-filled circular steel bridge piers”. ASCE

Journal of Bridge Engineering, Vol. 9 #1, Jan, pp 24-34. 8. Goodnight, J.C, Fulmer, S.J., Brown, N.R, Kowalsky, M.J., and Nau, J.M. (2014). “Experiential Applications of

3D Position Monitoring Systems for Earthquake Engineering Research” 14th National Conference on Earthquake Engineering, Anchorage, AK.

9. Priestley, M.J.N., Calvi, G.M., and Kowalsky, M.J. (2007) “Direct Displacement-Based Seismic Design of Structures” IUSS Press, Pavia Italy, ISBN 978-88-6198-000-6. 740 pages.

10. Jacobsen, L. S., “Damping in composite structures”, 2nd WCEE, Vol. 2, Japan, 1960, pp. 1029–1044.

D/t=35D/t=75D/t=155D.t=115

D/t=195

0%

5%

10%

15%

20%

1 2 3 4 5 6 7

Prop

osed

Tot

al D

ampi

ng

Coef

ficie

nt

Displacement Ductility