Post-Fire Bond Between the Steel Tube and Concrete in Concrete-filled Steel Tubular Columns

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Journal of Constructional Steel Research 67 (2011) 484–496

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Journal of Constructional Steel Research

journal homepage: www.elsevier.com/locate/jcsr

Post-fire bond between the steel tube and concrete in concrete-filled steeltubular columns

Zhong Tao a,b,∗, Lin-Hai Han c, Brian Uy a, Xian Chen b

a School of Engineering, University of Western Sydney, Penrith South DC, NSW 1797, Australiab College of Civil Engineering, Fuzhou University, Fuzhou, Fujian Province, 350108, Chinac Department of Civil Engineering, Tsinghua University, Beijing, 100084, China

a r t i c l e i n f o

Article history:

Received 28 June 2010

Accepted 17 September 2010

Keywords:

Concrete-filled steel tubes

Bond strength

Slip

Push-out

Composite action

Post-fire

a b s t r a c t

Push-out tests were carried out on 64 concrete-filled steel tubular columns, which had been exposedto ISO 834 standard fire for 90 min or 180 min, respectively. At the same time, 12 unheated specimenswere also prepared and tested for comparison. The variables investigated in the bond tests were selectedas (a) fire exposure time; (b) cross-section type; (c) cross-sectional dimension; (d) interface length todiameter (orwidth) ratio;(e) concrete type; (f) fly ash type; and(g) concrete curingcondition. Theeffectsof the above different parameters on the bond behaviour are discussed. The test results indicate thatfire exposure had a significant effect on the bond between a steel tube and its concrete core. A decreasein bond strength was generally observed for specimens after a fire exposure of 90 min; however, bondstrength recovery was found when the fire exposure time was extended to 180 min. Other factors alsohad influence to some extent.

© 2010 Elsevier Ltd. All rights reserved.

1. Introduction

Composite columns integrate the favorable characteristics of steel and concrete materials, therefore providing the merits of high strength and stiffness, high speed of construction, largeenergy dissipation, and considerable economy [1]. For this reason,concrete-filled steel tubular (CFST)columns have been used widelyin many countries in recent decades.

Nowadays, numerous research studies have been conductedon the performance of CFST members under different conditions,and some attention has been paid to the fire resistance of CFSTmembers or their residual properties after exposure to high tem-peratures [2,3]. Generally, the exposure of CFST members to firecan lead to significant loss of strength and stiffness. Researchalso showed that high temperatures had significant influenceon the residual bond strength between steel bars and concrete[4]. Furthermore, Hunaiti [5] tested 21 battened compositecolumns, which had been heat-treated at different temperaturesranging from 50 to 600 °C, to investigate the bond between steeland concrete. It was found that the loss in bond strength wassignificantly higher for high temperatures. A specimen tested at

∗ Corresponding author at: School of Engineering, University of Western Sydney,

Penrith SouthDC, NSW1797, Australia. Tel.:+61 2 47360064;fax:+61 2 47360137.

E-mail addresses: [email protected], [email protected] (Z. Tao).

600 °C showed a strength loss of about 83% when compared withits companion specimen tested at 100 °C. According to Chiang and

Tsai [4], bond degradation after fire exposure is mainly owing tothe fact that high temperatures cause water migration, dehydra-

tion of concrete,and decomposition of cement hydration products.Internal stress and thus micro- and macrocracks are generated

due to the inhomogeneous volume dilations of ingredients and thebuildup of vapor pressure in the pores.

It is expected that elevated temperatures will also have a signif-

icant impact on the bond between the steel tube and concrete inCFST columns. Although many studies have been performed in the

past to investigate the bond behaviour at ambient temperatures

for CFST columns [6–12], no research has been conducted so farconcerning their post-fire bond behaviour. An assessment of thedegree of deterioration of the bond in CFST columns after expo-

sure to fire can help engineers to evaluate the reliability of the loadtransfer between the steel tube and concrete at various locations if

necessary.

In this paper, a test program is introduced to investigate

the bond between the steel tube and concrete in CFST columnsthat have been exposed to and damaged by standard ISO 834

fire [13]. Push-out tests were further carried out to investigatethe influence of different parameters on the bond strength and

bond–slip curves. The bond mechanisms are also discussed in thispaper.

0143-974X/$ – see front matter © 2010 Elsevier Ltd. All rights reserved.doi:10.1016/j.jcsr.2010.09.006

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Z. Tao et al. / Journal of Constructional Steel Research 67 (2011) 484–496 485

Nomenclature

B Width of the square steel tubeD Diameter of the circular steel tubeEc Modulus of elasticity of concrete

f cu Compressive cube strength of concrete at the timeof the bond test

f cu, f Compressive cube strength of concrete at the timeof fire exposure

f y Yield strength of steelL Length of the steel tubeLi Length of the steel–concrete interfaceN max Maximum push-out loadN u Ultimate push-out load

p Interaction stress between the steel tube andconcrete

S SlipSu Slip corresponding to ultimate push-out load N ut Wall thickness of the steel tubeT Fire exposure timeτ Bond stress

τ cr Critical bond stressτ Rd Design bond strengthτ u Ultimate bond strength

2. Experimental program

2.1. General

A total of 76 specimens were prepared andtested. The variablesinvestigated were selected as (a) fire exposure time (0, 90 minand 180 min); (b) cross-section type (circular and square); (c)cross-sectional dimension (194–400 mm); (d) interface length todiameter or width ratio (Li/D or Li/B = 3, 5, or 7, where Li is the

length of the steel–concrete interface,D is thediameter of a circulartube, and B is thewidthof a squaretube); (e) concrete type (normalconcrete and self-consolidating concrete); (f) fly ash type (ClassII and Class III); and (g) concrete curing condition (sealed andunsealed before fire exposure).

Table 1 provides a summary for all circular and squarespecimens, in which specimen designations starting with C and Srefer to circular and square specimens, respectively. The followingnumber 1 to 9 in the labels is used to identify specimens withdifferent parameters. For clarity, the fire exposure time of 0, 90 or180 min for a specimen is also represented in its label. To obtainreliable test results, identical pairs were prepared and tested inmost cases, which are designated with suffixes ‘‘a’’ and ‘‘b’’ in thelabels, respectively.

2.2. Material properties

All tubes were manufactured from mild steel sheets. Since thecross-section size wasone of themajorvariables to be investigated,six kinds of steel sheets with different thicknesses were used in thefabrication to ensure that all tubes in a same test series of cross-section type had close cross-sectional aspect ratios (D/t or B/t ,where t is the wall thickness). A value of the B/t ratio of 40 waschosen for the square tubes, while the D/t ratios for the circulartubes were 37.3and 46.5, respectively. Therefore, all cross-sectionschosen were quite compact with an aim to assure that ultimatebond strength could be achieved before possible local bucklingoccurred for the steel tubes. Standard tensile coupon tests wereconducted to measure the material properties, where the average

measured values are given in Table 2. The obtained yield strengthsof steel ( f y) are also presented in Table 1.

Nowadays, there is an increasing interest in using self-consolidating concrete (SCC) to construct CFST columns. Known asa kind of high-performance concrete, SCC can flow in its gravityandfill in formworks without the need of anyinternal andexternalvibration. The utilisation of SCC in construction can shorten con-struction periods, reduce labour cost, improve compaction qualityespecially in confined and restricted zones where compaction isdifficult. In this test program, four mixes of SCC and a mix of nor-mal concrete (NC) were used to fill the tubes. The mix proportionsof the concretes are presented in Table 3. Well-graded limestonecoarse aggregate with a maximum size of 20 mm was used. To in-crease the slump and make the SCC more workable, fly ash (FA)and water reducer were added in the concrete mixes. Class II andClass III fly ashes in accordance with Chinese Standard GB/T 1596-2005 [14] were used to make the concrete, where Class II FA wasused for the concrete mixes of SCC-III and SCC-IV, and Class III FAwas used for those of SCC-I and SCC-II. According to GB/T 1596-2005,the type of FA is mainly classified in terms of fineness. Class IIFA is finer than Class III FA; therefore, the former is generally moreactive in affecting the concrete properties.

The measured material properties at room temperature andfresh properties of the SCC mixtures are given in Table 3, where

f cu, f and f cu are the compressive cube strengths of concrete at thetime of fire exposure and bond tests, respectively, and Ec is themodulus of elasticity at thetime of bond tests.Obviously, Class II FAconcrete had higher compressive strength compared with Class IIIFA concrete when the water-binder ratios and cement replacementratios were the same. It should be noted that the values of f cu, f aresmaller than those of corresponding f cu owing to a 7 month longercuring period of concrete when conducting bond tests.

2.3. Specimen preparation

Cold-formed steel tubes were used in the construction of circular specimens, while each square tube was fabricated bywelding four plates using a full penetration groove weld [15]. The

tubes were as received from themill andno specific measures weretaken to treat theinnersurfaces. Theoverall length(L)ofatubewasdesigned to be 50 mm longer than its length of the steel–concreteinterface (Li). The air gap of 50 mm located at the bottom of acolumn allowed the slip between the steel tube and concrete.Wooden plugs were fitted at the bottom of the specimens tosupport the wet concrete during casting. The plugs were removedlater to allow for travel of the concrete in the bond tests. Thesteel tubes were held upright and were filled with concreteflush to the top. To investigate the influence of concrete curingconditions, eight specimens (Series C7 andS7) presentedin Table 1were sealed immediately after the concrete was filled. This wasachieved by sealing their ends with polyethylene and wrapping anadditional thick plastic sheet to avoid possible moisture loss. Other

specimens were not sealed and just kept in indoor environment tocure.

2.4. Heat treatment

After being cured for about 4 months, the correspondingspecimens were heated by exposing them to heat in a furnaceat Structural Fire Resistance Laboratory in Tianjin Fire ResearchInstitute, China. Prior to the fire exposure, both ends of thespecimens were covered by a layer of mineral wool. This wasto reduce the heat transfer from the specimen ends to thesteel–concrete interface. The furnace heating was controlled asclosely as possible to the ISO 834 standard fire curve. The fireduration times (T ) weresetto be90 and 180 min,leading toa target

furnace temperature of 1006°

C and 1110°

C at the end of the firetests, respectively.



486 Z. Tao et al. / Journal of Constructional Steel Research 67 (2011) 484–496

Table 1

Specimen labels and test results.

No. Specimen label D × t × Li (mm) Li/D f cu (MPa) f y (MPa) T (min) N u (kN) N max (kN) τ u (MPa) Su (mm) Curve type Remarks

1 C1-0a 194 × 5.5 × 582 3 52.3 380 0 567.0 – 1.69 0.93 A

2 C1-0b 194 × 5.5 × 582 3 52.3 380 0 634.5 – 1.90 1.64 A

3 C1-90a 194 × 5.5 × 582 3 52.3 380 90 199.0 249.4 0.59 1.05 B

4 C1-90b 194 × 5.5 × 582 3 52.3 380 90 215.8 – 0.64 1.28 B

5 C1-180a 194 × 5.5 × 582 3 52.3 380 180 500.0 – 1.49 3.72 A6 C1-180b 194 × 5.5 × 582 3 52.3 380 180 448.1 – 1.34 3.45 A

7 C2-0 194 × 5.5 × 582 3 72.1 340 0 372.3 – 1.11 1.26 B

Class II fly ash used8 C2-90 194 × 5.5 × 582 3 72.1 340 90 122.6 – 0.37 0.72 B

9 C2-180 194 × 5.5 × 582 3 72.1 340 180 654.2 – 1.96 3.55 B

10 C3-0 194 × 5.5 × 582 3 53.2 340 0 929.7 – 2.78 2.66 A

Class II fly ash used11 C3-90 194 × 5.5 × 582 3 53.2 340 90 344.5 – 1.03 1.76 B

12 C3-180 194 × 5.5 × 582 3 53.2 340 180 673.2 – 2.01 3.14 A

13 C4-0a 377×8.1×1131 3 52.3 363 0 612.3 754.2 0.48 1.41 B

14 C4-0b 377×8.1×1131 3 52.3 363 0 549.7 941.3 0.43 1.34 B

15 C4-90a 377×8.1×1131 3 52.3 363 90 978.8 – 0.76 3.42 B

16 C4-90b 377×8.1×1131 3 52.3 363 90 886.5 1072.0 0.69 1.98 B

17 C4-180a 377×8.1×1131 3 52.3 363 180 1465.0 – 1.14 2.22 B18 C4-180b 377×8.1×1131 3 52.3 363 180 1239.3 – 0.97 2.33 B

19 C5-90a 194 × 5.5 × 582 3 46.1 380 90 241.0 – 0.72 1.23 A

20 C5-90b 194 × 5.5 × 582 3 46.1 380 90 379.4 – 1.13 2.85 A

21 C5-180a 194×

5.5×

582 3 46.1 380 180 515.2 – 1.54 6.15 A22 C5-180b 194 × 5.5 × 582 3 46.1 380 180 567.7 – 1.70 6.66 A

23 C6-90a 194 × 5.5 × 582 3 63.3 380 90 181.0 – 0.54 0.56 B

Normal concrete24 C6-90b 194 × 5.5 × 582 3 63.3 380 90 288.4 – 0.86 1.34 B25 C6-180a 194 × 5.5 × 582 3 63.3 380 180 578.4 – 1.73 2.96 B

26 C6-180b 194 × 5.5 × 582 3 63.3 380 180 562.3 – 1.68 3.21 A

27 C7-90a 194 × 5.5 × 582 3 52.3 380 90 170.2 – 0.51 0.78 B

Concrete cured in

sealed condition

28 C7-90b 194 × 5.5 × 582 3 52.3 380 90 437.1 – 1.31 3.88 B

29 C7-180a 194 × 5.5 × 582 3 52.3 380 180 549.0 – 1.64 3.57 A

30 C7-180b 194 × 5.5 × 582 3 52.3 380 180 519.7 – 1.55 4.00 A

31 C8-90a 194 × 5.5 × 970 5 52.3 380 90 522.6 – 0.94 2.35 A

32 C8-90b 194 × 5.5 × 970 5 52.3 380 90 602.6 – 1.08 4.29 B

33 C8-180a 194 × 5.5 × 970 5 52.3 380 180 594.2 – 1.07 6.40 B

34 C8-180b 194 × 5.5 × 970 5 52.3 380 180 595.8 – 1.07 2.78 B

35 C9-90a 194×5.5×1350 7 52.3 380 90 611.6 – 0.79 2.62 B

36 C9-90b 194×5.5×1350 7 52.3 380 90 545.8 – 0.70 1.41 B

37 C9-180a 194×5.5×1350 7 52.3 380 180 740.6 – 0.95 2.72 B

38 C9-180b 194×5.5×1350 7 52.3 380 180 651.9 – 0.84 2.94 B39 S1-0a 200 × 5 × 600 3 52.3 320 0 140.3 225.4 0.31 0.52 B40 S1-0b 200 × 5 × 600 3 52.3 320 0 195.6 324.5 0.43 0.74 B

41 S1-90a 200 × 5 × 600 3 52.3 320 90 126.1 – 0.28 2.36 A

42 S1-90b 200 × 5 × 600 3 52.3 320 90 95.2 170.0 0.21 2.69 C

43 S1-180a 200 × 5 × 600 3 52.3 320 180 173.2 – 0.38 2.52 A

44 S1-180b 200 × 5 × 600 3 52.3 320 180 136.5 – 0.30 1.52 B

45 S2-0 200 × 5 × 600 3 72.1 297 0 165.8 299.0 0.36 0.58 B

Class II fly ash used46 S2-90 200 × 5 × 600 3 72.1 297 90 164.8 215.2 0.36 2.50 B47 S2-180 200 × 5 × 600 3 72.1 297 180 347.1 – 0.76 2.87 A

48 S3-0 200 × 5 × 600 3 53.2 297 0 148.7 272.9 0.33 0.68 B

Class II fly ash used49 S3-90 200 × 5 × 600 3 53.2 297 90 128.4 166.1 0.28 2.28 B

50 S3-180 200 × 5 × 600 3 53.2 297 180 349.0 – 0.77 4.47 A

51 S4-0a 400× 10× 1200 3 52.3 427 0 269.7 512.6 0.15 1.08 B

52 S4-0b 400× 10× 1200 3 52.3 427 0 308.2 587.4 0.17 1.05 B

53 S4-90a 400× 10× 1200 3 52.3 427 90 194.1 350.8 0.11 2.68 C

54 S4-90b 400× 10× 1200 3 52.3 427 90 305.8 510.6 0.17 1.10 C55 S4-180a 400× 10× 1200 3 52.3 427 180 403.5 570.3 0.22 1.32 C

56 S4-180b 400× 10× 1200 3 52.3 427 180 407.4 580.0 0.22 2.83 B

57 S5-90a 200 × 5 × 600 3 46.1 320 90 76.8 98.1 0.17 2.92 B

58 S5-90b 200 × 5 × 600 3 46.1 320 90 79.4 102.6 0.17 3.97 C

59 S5-180a 200 × 5 × 600 3 46.1 320 180 71.6 – 0.16 1.88 A

60 S5-180b 200 × 5 × 600 3 46.1 320 180 78.7 102.9 0.17 1.77 C

61 S6-90a 200 × 5 × 600 3 63.3 320 90 121.6 141.6 0.27 2.83 C

Normal concrete62 S6-90b 200 × 5 × 600 3 63.3 320 90 135.5 141.9 0.30 1.68 B

63 S6-180a 200 × 5 × 600 3 63.3 320 180 230.3 – 0.51 2.88 A

64 S6-180b 200 × 5 × 600 3 63.3 320 180 252.3 – 0.55 2.75 A

65 S7-90a 200 × 5 × 600 3 52.3 320 90 78.7 152.6 0.17 1.73 C

Concrete cured in

sealed condition

66 S7-90b 200 × 5 × 600 3 52.3 320 90 56.2 110.0 0.12 3.30 C

67 S7-180a 200 × 5 × 600 3 52.3 320 180 250.6 – 0.55 2.89 A

68 S7-180b 200 × 5 × 600 3 52.3 320 180 211.3 – 0.46 1.92 B

69 S8-90a 200 × 5 × 1000 5 52.3 320 90 202.9 356.4 0.27 2.81 C

70 S8-90b 200×

5×

1000 5 52.3 320 90 102.2 354.2 0.13 0.96 C




Table 1 (continued)

No. Specimen label D × t × Li (mm) Li/D f cu (MPa) f y (MPa) T (min) N u (kN) N max (kN) τ u (MPa) Su (mm) Curve type Remarks

71 S8-180a 200 × 5 × 1000 5 52.3 320 180 419.7 482.3 0.55 2.41 C

72 S8-180b 200 × 5 × 1000 5 52.3 320 180 400.0 554.2 0.53 2.84 C

73 S9-90a 200 × 5 × 1400 7 52.3 320 90 420.3 691.0 0.40 3.27 C

74 S9-90b 200 × 5 × 1400 7 52.3 320 90 365.8 594.5 0.34 3.07 C

75 S9-180a 200 × 5 × 1400 7 52.3 320 180 610.6 – 0.57 – –

76 S9-180b 200 × 5 × 1400 7 52.3 320 180 609.7 – 0.57 – –

Table 2

Material properties of steel.

Typ e Thickness

(mm)

Elastic modulus

(kN/mm2)

Yield strength

(N/mm2)

Poisson’s

ratio

Ultimate strength

(N/mm2)

Elongation

percentage (%)

Used in specimen

series

Type I 5 197 320 0.281 437 41.7 S1, S5–S9

Type II 5 195 297 0.276 368 42.5 S2, S3

Type III 5.5 198 380 0.269 433 28.8 C1, C5–C9

Type IV 5.5 201 340 0.270 460 35.4 C2, C3

Type V 8.1 187 363 0.282 433 28.5 C4Type VI 10 205 427 0.265 383 24.5 S4

Table 3

Mixture and material properties of concrete.

Type Water(kg/m3)

Cement(kg/m3)

Fly ash(kg/m3)

Sand(kg/m3)

Aggregate(kg/m3)

Water reducer(kg/m3)

Slump(mm)

Slump flow(mm)

f cu, f f cu Ec

(N/mm2)Used in specimenseries

SCC-I 190 260 240 740 950 5.4 265 640 25.3 46.1 27 300 C5, S5SCC-II 170 380 170 800 835 6.08 240 560 45.3 52.3 29 500 C1, C4, C7-C9, S1,

S4, S7-S9

SCC-III 190 260 240 740 950 5.4 255 650 40.0 53.2 35 500 C3, S3

SCC-IV 170 380 170 800 835 6.08 260 590 57.3 72.1 38 900 C2, S2

NC 210 552 – 611 1137 – – – 50.4 63.3 34 500 C6, S6

In a real fire situation, thermal creep deformation will occur forthe steel and concrete materials in a loaded column. Owing to thelimitation of the furnace facility, the specimens were not loadedduring the fire exposure in the current test program. Furtherresearch is required in the future to address this issue.

2.5. Test setup and instrumentation

Bond tests were conducted about 7 months later after the heattreatment. Prior to the bond tests, the top surfaces of all specimenswere ground smooth to provide a level surface and to ensureeven distribution of the applied load. A 2000 kN capacity testingmachine was used to conduct the bond tests as shown in Fig. 1. Thespecimen was set up in the testing machine in a vertical positionwith the air gap at the bottom. A steel block was placed betweenthe specimen and the top platen of the testing machine. The steelblock had a cross-section which was a little smaller than that of the concrete core. This assured the load to be applied only onthe concrete core and allowed the concrete core to be pushed

downward when testing. The load was measured using a load cellwhich was placed on the steel block.All specimens were tested under monotonic loading with a

loading rate of about 0.5 mm/min. Axial strains in the steel weremeasured using strain gauges installed along the length of steeltubes. The spacing between two adjacent strain gauges rangedfrom 125 to 150 mm. It is possible to analyse the load transfermechanism from the concrete core to the steel tube based on thestrain readings. In the current tests, both slip values at the loadedand free ends were measured. This was achieved by readings of seven linear variable displacement transducers (LVDT) mountedon the specimen as shown in Fig. 1. LVDT 1 and LVDT 2 were usedto measure the slip at the loaded end, while LVDT 3 to LVDT 6 wereused to monitor initial settlements of the steel tube at the bottom.Therefore, the free end slip was calculated from the readings of LVDT 7 by subtracting the obtained initial settlements.

3. Test results and discussion

A general view of the specimens after testing is shown inFig. 2. The failure mode was characterised by the slip between thesteel tube and concrete, which could be observed on the top of a

specimen. There was no other visible change in appearance for thesteel tubes exceptthe spalling offof oxide layers formedduring thefire exposure.

3.1. Bond stress (τ ) versus slip (S) curves

Average bond stress (τ ) is used in this paper to evaluate thebond behaviour between a steel tube and its concrete core, whichis the push load divided by the area of the interface. Therefore, τ in this paper refers to the average bond stress unless otherwiseindicated. Although the distribution of local bond stress along atube and across its perimeter may not be uniform, the concept of average bond stress is quite straightforwardand can be interpretedand used by engineers very easily.

Fig. 3 shows typical average bond stress (τ ) versus loaded-endslip (S) curves. For clarity, a slip in this paper refers to the onemeasured at the loaded end in the following unless otherwisenoted. Generally, the obtained τ –S curves can be classified intothree types, which depend on the mutual influences of differentparameters. The corresponding curve types for all circular andsquare specimens are summarised in Table 1. It should be notedthat no curve type is assigned to the specimens S9-180a and S9-180b. This is owing to the fact that these two square specimenswere the longest and had been exposed to fire for 180 min. Dueto the fire damage to the steel, the tubes buckled unexpectedly atthe end of the bottom before the ultimate bond strength could bedeveloped between the steel tube and concrete. Without concreteat the bottomto support the steel tube, a pre-mature weld fracture

occurred at the end of testing. Thus, particular attention should bepaid in any further bond tests to avoid this problem.




a b

Fig. 1. Push-out test setup.

(a) All specimens after testing. (b) Typical circular section. (c) Typical square section.

Fig. 2. A general view of specimens after testing.

(a) Type A and type B. (b) Type C.

Fig. 3. Typical bond stress (τ ) versus slip (S) curves.

Type A curve shown in Fig. 3(a) is the typical τ –S relationshipwith a falling branch beyond the maximum bond stress. In thecurrent tests, about one-third of circular specimens and one-fifthsquare specimens demonstrated this feature. The majority of thespecimens are those without fire exposure or with a fire exposureof 180 min. For a Type A curve, it has an initial linear portion(OA) up to 50%–80% of the peak bond strength, followed by atransitional portion (AB). After reaching the peak strength (PointB), a rapidly declining portion BC appears before relatively stableresidual strength can be achieved. Although the residual strengthkeeps decreasing from Point C to Point D, the curve tends to flattenas it approaches Point D. In some cases, the curve can be a nearlyhorizontal flat one after reaching Point C. The residual strength atthe end of the testing is about 35%–65% of the peak strength for

Type A circular specimens, while that for the square specimensranges from 60% to 85% of the peak strength.

The τ –S curves of about two-thirds of circular specimens andtwo-fifths of square specimens can be classified into Type B group.Compared with Type A curves, a Type B curve shows quite similarfeatures except the existence of the second ascending portion C′D′.At Point C′, the minimum residual strength is attained, whichis about 40%–80% of the bond strength at Point B ′ for circularcolumns, and 50%–90% of that for square columns. The bond stressincreases once again beyond Point C′. Compared with the residualstrength at Point C′, the bond stress at the end of testing (Point D ′)generally increases by 20%–120%. Therefore, of the Type B curves,the residual strengths of three circular specimens and the majorityof square columns are even much higher than their correspondingbond stresses at Point B′. The strength increase ranges from 20%to 80%. The maximum push-out loads (N max) recorded at the end

of testing are presented in Table 1 if they are larger than theircorresponding ultimate loads.




(a) Circular specimen (C9-90b). (b) Square specimen (S9-9 0b).

Fig. 4. Axial strain variations along length of tubes.

A typical Type C τ –S curve is depicted in Fig. 3(b), whichis characterised by the lack of any falling branch. For the testspresented in this paper, Type C curves are only observed forabout two-fifths of square specimens, and most of them had beenexposed to fire for 90 min. No circular specimens can be classified

into this group. Initially, there is an elastic response until reachingPoint A′′, where the bond stress is about 60%–80% of the ultimatebond strength at Point B′′, followed by the inelastic response withgraduallydecreasing slope. Afterreaching the transitional PointB ′′,the τ –S response is almost linear up to the end of the testing (PointC′′). Generally, the bond stress at Point C′′ is 1.3–1.9 times the bondstrength at Point B′′.

3.2. Bond mechanisms and bond stress distribution

To understand the different responses in terms of τ –S rela-tionship, it is necessary to check the bond mechanisms betweenthe steel tube and concrete first. It is well known that bond canbe considered as a result of three different mechanisms: chemi-

cal adhesion, friction and macro-interlocking [6,16]. The friction,which was also called as micro-interlocking by some researchers[6], mainly arises from the roughness of the steel surface, whilethe macro-interlocking is owing to the surface irregularities re-lated to the manufacturing tolerances. Both friction and macro-interlocking can cause mechanical keying of the concrete core withthe steeltube, and therefore macro-interlocking can alsobe viewedas friction. From another point of view, the macro-interlocking,somehow, is comparable to the mechanical interlocking betweenthe ribs of deformed reinforcement bars and concrete.

In practice, it is very difficult to separate the three componentsof bond since the distribution of surface irregularities is quite vari-able and uneven. Moreover, there is no practical criterion that canbe used to distinguish between the micro-interlocking and macro-

interlocking. To make matters more complicated, the shrinkageof concrete has significant influence on the bond behaviour andmay affect the bond contribution of different components. How-ever, some efforts were made by Virdi and Dowling [6], wherethree circular specimens of Group ST had their inside surfacesmachined to a smooth finish, and another group of three circu-lar specimens (Group PH) had their inner surfaces covered with alubricant in five coats. Therefore, the surface roughness of GroupST was greatly diminished, and the bond strength was believedto have been resulted mainly from the chemical adhesion. On theother hand, the bond for specimens in Group PH was largely con-tributed by the macro-interlocking effect. The test results of GroupST and PH were further compared with those of normal specimens(Group CS) without any surface treatment. It seems that the bond

strength from chemical adhesion only (Group ST) could attain avalue of about 19% ultimate bond strength in normal conditions,

butit failedand decayed rapidly. Theslip corresponding to thepeakbond strength was about 0.5 mm for Group ST. The response of Group PH showed a near absence of an initially stiff region, but thebond resistance resulted from macro-interlocking kept increasinguntil the end of testing. After compared with the tests results of

Group CS, Virdi and Dowling [6] postulated that the ultimate bondstrength for CFST columns was mainly contributed by friction.

To further understand the bond mechanisms, it is necessary tocheck the bond stress distribution along the length of tubes. Thebond stresses can be obtained based on the axial strain gradientsin the steel tube [8]. According to Roeder et al. [8], the axialstrain gradients are related by statics to the interface bond stressdistribution. If N x is the axial load in the steel tube at location x,then the bond stress, τ( x), can be given by

τ( x) =dN x

C · d x(1)

where C is theperimeter of the concrete section in contact with thesteel tube, and is equal to π(D − 2t ) and 4(B − 2t ), respectively,

for circular and square sections.Figs. 4 and 5 demonstrate the axial strain variations and bond

stress distributions along the length of two typical specimens,respectively. It should be noted that the axial strains measuredat the bottom were normally smaller than the expected values.This is owing to the influence of the constraint conditions at thebottom as shown in Fig. 6, where the steel tube had a tendency todeform outwardly due to the Poisson’s effect, but was restrainedat the end. Therefore, the measured strains near the bottom werenotused to calculate the bond stresses at that location,which werealternatively determined based on theassumption that allthe pushload was carried by the steel tube alone at the bottom.

Apparently, the trends of the axial strain distributions shownin Fig. 4 for the circular and square specimens are quite different.

For the circular specimen C9-90b, the axial strain of the steeltube increased very fast within a short distance away from theloaded end. It indicates the fast development of local bond stressbetween the steel tube and concrete in this area. Since only theconcrete wasloaded on the top, it would dilateunder compression,and generated radial contact pressure ( p) between the steel tubeand concrete, as shown in Fig. 6. The radial pressure resultedfrom the confinement of the steel tube could further enhance thefrictional resistance and therefore bond stress at the top, as shownin Fig. 5(a). In other words, strong composite action was formedbetween the circular steel tube and concrete. As the majority of the load was transferred from the concrete to the steel tube atthe top, the compression was low in the concrete and high in thesteel in the lower part of the column. Therefore, small or even

no contact pressure would be formed between the steel tube andconcrete there. This led to the much slower increase of axial strain




(a) Ci rcular spec imen (C9 -90b). (b) Square specimen (S9-90b).

Fig. 5. Bond stress distributions at different load levels.

in the lower part of the tube, and the bond stresses in this areawere also much smaller than those at the top. As far as the squarespecimen S9-90b is concerned, the axial strain distribution alongthe length was close to a simple triangle as shown in Fig. 4(b).This demonstrates that the square column was not so efficient

in transferring load between the steel tube and concrete as itscircular counterpart. Friction could be formed uniformly aroundthe perimeter of a circular tube, but mainly formed only in thevicinity of the corners of a square tube [7]. Therefore, the squarecolumn S9-90 took a much longer length to develop reliable bond,andthe bond stresses,in general, distributed more evenlyalongthelength of the specimen. The fluctuation in the stress distributionshown in Fig. 5(b) was attributed to the strain measurement errorsand especially the irregularities of the steel tube. It is worth notingthat the bond stresses at the bottom were often larger than thoseat its adjacent measure points. This is somewhat owing to thepinching effect as shown in Fig. 6. But this effect was not expectedto be critical since the steel tube remained elastic during the tests.Another important reason is that there was local slip and bondstress transfer occurred at the bottom since the total load of the

specimen was resisted by the steel.In the current test program, the free-end slip was also

measured. A typical bond stress versus free-end slip curve forspecimen C4-0a is shown in Fig. 7. Therefore, the initiation of theslip at the free end can be recognised. Generally, very small slipat the free end occurs in the initial loading stage. The slip at thefree end develops much more quickly after Point A is reached,which approximately coincides with the termination of the elasticresponses of τ –S curves shown in Fig. 3. Therefore, during theelastic stage,the bond resistance is contributed mainly by chemicaladhesion and friction. In most cases, the chemical adhesion isonly active in the initial loading stage since the average loaded-end slip at the end of elastic stage for all the test specimens is1.65 mm in the current tests. This value is much larger than the

value of 0.5 mm corresponding to the peak bond strength forGroup ST reported by Virdi and Dowling [6]. Although the maro-interlocking mechanism also starts to work in this stage afterthe slip initiates at the top, it is not significant according to thetest results presented by Virdi and Dowling [6]. After the notableslippage occurred at the free end, a second stage of behaviourdevelops characterised by a gradual softening of the stiffness of τ –S response. The reason is that the adhesion is now negligible,and the bond capacity coming from the friction begins to reduce aswell. The friction reduction could be attributed to the shearing off of mortar which extends into small pits in the steel tube. Now, themacro-interlocking starts to play an important role in this stage.After the ultimate bond strength has been achieved, a third stageof behaviour develops, during which the friction is diminishing

rapidly in most cases while the bond resistance from macro-interlocking keeps increasing. For specimens with high ultimate

Fig. 6. Deformation of steel tube under push-out force.

bond strength, the effect of diminishing friction will outweigh thebeneficial influence of the increasing macro-interlocking. In thiscase, the τ –S response will be a Type A or Type B curve with anobvious declining branch. Otherwise, it will show the feature of Type C curves. Actually, the transition between the second andthe third stages was not always as smooth as that shown in Fig. 3.Sometimes, sudden decrease in bond strength was observed afterthe ultimate bond strength had been reached, as shown in Fig. 8.But part of the bond strength recovered soon. This phenomenonwas also observed and reported by some other researchers [9,10].In this paper, three circular specimens (C1-90a, C1-90b, and C4-90b) and almost all the unheated square specimens demonstratedthis feature. Forthesespecimens, it seems theadhesionand friction

reduced too fast before the macro-interlocking could absorb theenergy stored in the specimens. All these specimens are classifiedinto Type B group in this paper. As mentioned earlier, the onlydifference between Type A and Type B curves is the existence of a second ascending portion. For a Type B curve, it seems that theresidual strength is mainly contributed by the macro-interlocking,thus the residual strength increases during the later loading stage.This might happen if the residual bond resistance contributed byfriction is minor or the macro-interlocking grows very fast forspecimens with large fabrication tolerances.

Identical specimens have been prepared to conduct push-outtests for most test series. Generally, the agreement is reasonablefor each pair of duplicate specimens in terms of ultimate bondstrength and τ –S curve. Therefore, bond strengths obtained from

the average of two individual push-out results are used andcompared in the following discussion. It is worth noting that,




(a) Full-range τ –S curves. (b) Initial τ –S curves.

Fig. 7. Comparison of slip at loaded end and free end (C4-0a).

Fig. 8. Sudden load reduction after peak loads were reached for some specimens.

Fig. 9. Discrepancy in τ –S curves for identical specimens.

in some cases, a notable discrepancy can be spotted, which isconsistentwith the observation presented by otherresearchers [8].A typical example is shown in Fig. 9, where the τ –S curve of specimen S7-180a can be classified into Type A while that of specimen S7-180b was recognised as Type B. Moreover, there isa difference of 18.6% in the measured peak bond strength betweenthe two specimens, although the initial slopes are quite close.Obviously, this discrepancy canbe attributed to thevariation in the

surface roughness of the two specimens.

3.3. Bond strength (τ u)

Currently, there is no single definition for bond strength thatis universally accepted. Virdi and Dowling [6] conducted push-outtests on over 100 concrete-filled circular steel tubes, where themaximum stresses were achieved at the end of the tests when theconcrete core attained its travel limit. In other words, all curvesdemonstrated by Virdi and Dowling [6] can be identified as Type Ccurves, as shown in Fig.3(b).This observation is notconsistent withthat observed by others. The explanation is that the tubes usedby Virdi and Dowling [6] had very small diameter-to-thicknessratios ranged from 24 to 36. So they were highly resistant to

lateral expansion and therefore could resist high interface shearstress and the macro-interlocking were magnified as well. The

reported bond strengths ranged from 1.4 to 2.8 MPa, in general,are also higher than those reported by others. A conservativeestimation of bond strength was made by Virdi and Dowling [6]on the basis of a critical bond strain of 0.0035. Aly et al. [12]believed that thechemicalbond is unstable,and post-peakresidualbond stress was recommended to be taken as the bond strength.However, this method is only applicable to Type B curves in

defining reliable residual bond strength. For other types of curves,the determination of residual strength will be rather arbitrary. Thebond between thecircularsteeltubeand concrete wasinvestigatedby Roeder et al. [8], where bond stress capacity was defined asthe average interface stress associated with the initial rigid bodyslip of the concrete core relative to the steel tube. According toRoeder et al. [8], the behaviour of concrete was of rigid bodymotion between the core and the tube with reducing mechanicalresistance from interface shear after the slip was greater than thatat the peak load. For this reason, maximum average bond stresswas used by Shakir-Khalil [7],Roederet al. [8], Nezamianet al. [10],and Nardin and El Debs [11].

According to the above discussion, maximum average bondstress is also adopted in this paper to represent ultimate bondstrength (τ u) for Type A and Type B curves. The average values

of the corresponding slip (Su) for circular specimens with a fireexposure of 0 min, 90 min and 180 min are 1.54 mm, 1.97 mmand 3.74 mm, respectively, while those for square specimens are0.78 mm, 2.35 mm, and 2.64 mm, respectively. However, the bondstrength definition is notso straightforwardfor Type C curves sincethe bond stress keeps increasing during the whole loading process.In this paper, the bond stress at the transitional Point B ′′ shown inFig. 3(b) is defined as the ultimate bond strength. This is justifiedby the fact that the average value of Su is 2.45 mm for all TypeC curves, which is quite close to the average value of other fire-damaged square specimens.

The values of τ u and Su for all circular and square specimens aresummarised in Table 1. At the same time, the ultimate push-outloads (N u) corresponding to τ u are also given in this table.

3.4. Critical bond stress

The average stress corresponding to Point A on the bond stressversus free-end curve shown in Fig. 7 is designated as critical bondstress (τ cr) in this paper. As mentioned earlier, the termination of the elastic response of a τ –S curve normally corresponds to thefast development of the slip at the free end. Therefore, the criticalbond stress is one of theimportant characteristics of τ –S curves. Ascan be seen from Fig. 7, the slipat the free end is much smaller thanthat at the loaded endbefore τ cr is reached.Thisis owing tothe factthat the slip at the free end during this stage is mainly generatedfrom the compressive deformation of the steel tube at the bottom.But beyond Point A, it suggests the change in slip mechanisms

and indicates that the concrete core begins to slip along the wholeinterface.




(a) Circular specimens. (b) Square specimens.

Fig. 10. Normalized critical bond stress.

( a) I nn er s urf ace o f the cir cular tube. ( b) I nn er s urf ace o f the sq uare tube.

Fig. 11. Friction marks observed on steel tubes.

Fig. 10 shows the normalized τ cr with respect to the correspond-ing ultimate bond strength (τ u). Obviously, there is considerable

scatter in τ cr/τ u values across the sample. But τ cr/τ u values are, ingeneral, bigger than 0.5. For circular columns, smaller τ cr/τ u valueswere achieved for those fire-damaged specimens compared withthose undamaged ones. But for square columns, the trend is justthe opposite. The explanation could be that friction contributessignificantly to the bond strength for circular columns before τ cr

is reached, while for square columns, the friction is less effective.This is confirmed by the fact that friction marks were observedaround the perimeter of a circular tube, but mainly found in thevicinity of the corners of a square tube, as shown in Fig. 11. Thisphenomenon was also observed by Shakir-Khalil [7]. It looks likethat concrete strength has strong influence on friction accordingto its formation mechanism mentioned earlier. Long-time fire ex-posure will decrease the concrete strength, which in turn causesthe earlier shearing off of the bonding mortar and reduces thefriction between the steel tube and concrete. For this reason, thefire-damaged circular specimens attained their ultimate state soonafter τ cr was reached, and τ cr/τ u increased accordingly. But forthe square specimens, chemical adhesion was more important incontributing to the bond strength beforeτ cr was reached since fric-tion was less effective in this case. Not like friction, the whole in-terface can provide adhesion to resist the push-out load. This isproved by the fact that the inner surfaces of square steel tubeswere still covered with the remainder of mortar after the push-out testing, as shown in Fig. 11(b). Therefore, the square tubes hadlarger τ cr/τ u values at ambient temperatures since chemical adhe-sion was likelyto be conqueredjust beforethe rigid body slip of theconcrete occurred. But after exposure to fire, the chemical adhe-sion was affected by fire damage, leading to the earlier occurrence

of slip andsmallerτ cr/τ u. Itcan alsobe found from Fig. 10 thatthereis no statistically significantdifferencebetween theτ cr/τ u ratios for

fire-exposed circular and square specimens, and the fire exposureperiod has no significant influence on τ cr/τ u either.

3.5. Influence of different parameters

3.5.1. Effects of fire exposure time and cross-section type

It is no doubt that fire exposure has impact on the chemicaladhesion and friction between the steel tube and concrete.Furthermore, it can lead to the change of concrete volume, andthus affects the composite action between the steel tube andconcrete. The volume change of concrete causes a radial reductionor enlargement. Its influence on the bond behaviour has been welldocumented by Roeder et al. [8].

According to the test results reported by Khoury [17], after ini-tial expansion, cement pastes could undergo considerable shrink-age above 100°C, as they lost evaporable water andas dehydration

of the various hydrates occurred at different temperatures. Mean-while, substantial residual expansion was observed for aggregatesin concrete. The tests conducted by Khoury [17] indicate that lime-stone aggregate exhibited residual expansion after cooling fromabout 350 °C demonstrating thermal instability at that tempera-ture. A third contributing factor to the volume change of concreteis crack development. Micro-cracking formed during cooling canlead to significant residual expansion of concrete. Therefore, resid-ual strain of concrete after exposure to fire will greatly dependon the temperature history, aggregate type, cement type, water-binder ratio, and cement replacement ratio used. It is worth notingthat post-fire concrete may also experience some volume changeafter absorbing moisture from the atmosphere. However, this in-fluence is minimized for CFST columns since the concrete is cast inthe steel tubes.

At this moment, it is still very difficult to accurately predictthe residual strain of concrete after fire exposure. This is owing




Fig. 12. Residual strain of quartz and limestone concrete after cooling [18].

to the limited test results available and the substantial number of affecting parameters involved. Generally, residual shrinkage canbe expected for concrete subjected to lower temperatures whileresidual expansion occurs when exposed to higher temperatures(Fig. 12). According to Schneider [18], the first case occurred formost concretes after heating up to less than 400 °C, and thesecond case was observed by cooling of concretes being heated

above 400 °C. A residual dilation of up to 6000 µε was reportedby Schneider [18] for unloaded limestone concrete experienced aheating of 800 °C, as shown in Fig. 12.

Fig. 13 depicts the temperature distributions along the x axisfor specimens with different cross-sectional dimensions at theend of fire exposure. The temperatures were calculated based onheat transfer analysis by using the ABAQUS software [19]. Moredetails about the analysis can be found in [20]. Since concretehas a relatively low thermal conductivity, it is not surprisingthat the temperatures are much lower within the concrete corethan those near the concrete surface (Point 3 shown in Fig. 13),especially for these big columns. In Fig. 13, Point 1 is the centrepoint of the cross-section and Point 2 is the midpoint betweenPoint 1 and Point 3. The temperatures of these characteristic

points calculated are summarised in Table 4. It seems that sectiontype or cross-sectional dimension has no significant influence onthe temperatures at the interface. For columns with 90 min fireexposure, the calculated temperature at Point 3 ranges from 877to 925 °C, while that for columns with a T of 180 min ranges from1016 to 1050 °C. Since the temperature distribution across thewhole section is uneven, the temperature at Point 2 is chosen asthe characteristic temperature (T char) of the section to simplify theanalysis. Therefore, for the smaller circular and square columns,the values of T char are all larger than 500 °C. It is expected thatsubstantial residual expansion occurred for the concrete in thesecolumns. But for the bigger columns, all T char are under 400 °C

except that of the circular columns with a diameter of 377 mmand a fire exposure of 180 min. In this exceptional case, T char is501 °C. Hence,the biggercolumns arenot expected to have obviousvolume change according to Schneider [18].

Fig. 14 demonstrates the effect of fire exposure time (T ) on τ –S

curves. As can be seen, the ultimate bond strength τ u generallydecreased after exposure to fire for 90 min. For circular columns,

significant loss of bond strength occurred with a strength loss of 73–77%. The strength loss for square columns was comparativelymuch less, which ranged from 12% to 34%. The deterioration inbond strength was mainly owing to the detrimental influence of fire exposure on chemical adhesion and friction. The greater loss of bond strength for circular columns indicates that the fire exposurehas much more profound influence on friction. For specimenswith a fire exposure of 180 min, it is interesting to note that thebond strength began to reversely increase, rather than to declineany further. In some cases, the strength could be fully recovered,which was especially true for those square columns. Two mainreasons can be used to explain this. On the one hand, the valuesof T char for the smaller specimens were quite close to 900 °C,therefore, significant residual expansion would be expected. Onthe other hand, the more severe the concrete damaged, the easierthe concrete dilates under axial compression. The radial expansionof concrete could increase the contact pressure and thereforefriction resistance. Another possible beneficial effect resulted fromthe residual expansion is that it could counteract the influence of concrete shrinkage formed during the curing stage.

Fire exposure also affected the initial slope of τ –S curves, whichmainly came from chemical adhesion between the steel tube andconcrete. Generally, the stiffness degradation could be as high as50%. But it seems that the fire exposure time had no significantinfluence on the initial slope. This is owing to the fact that theinterface temperatures were quite close when the specimens wereexposed to fire either with a time of 90 or 180 min. Therefore, thedamage to the chemical bond were expected to be quite close.

Fig. 15 compares the bond strength of circular specimenswith those of square specimens. It is clear that circular columnsgenerally had much higher bond strength than square columns.This can be explained by the aforesaid different bond mechanismsfor circular and square columns, i.e., circular columns are muchmore effective in transferring the interface shear stress. The flatparts of a square tube are susceptible to bend outwardly, oncecontact pressure is formed between the steel tube and concrete.This has an adverse effect on the actual contact between the steeltubeand concrete. Therefore,the effective contact likelyforms onlyin the vicinity of the corners for a square tube, which can be called‘‘effective contact area’’.


Fig. 13. Temperature distribution along x axis.




Table 4

Calculated temperatures at different locations.

Section type Cross-sectional dimension (mm) T = 90 min T = 180 min

Point 1 Point 2 Point 3 Point 1 Point 2 Point 3

Circular194 489 620 925 886 933 1050

377 130 257 898 354 501 1029

Square200 429 531 913 803 867 1044

400 76 178 877 251 393 1016

S3

S3

S3


Fig. 14. Effect of fire exposure time on τ –S curves.


Fig. 15. Effect of cross-section type on bond strength.


Fig. 16. Effect of cross-sectional dimension on bond strength.

3.5.2. Effect of cross-sectional dimension

The effect of cross-sectional dimension on τ u is depicted inFig. 16. In general, the bond strength decreased significantly whenD or B increased. This is consistent with the observation madeby Roeder et al. [8]. The larger the tube size, the bigger the gapforms between the steel tube and concrete due to the shrinkageof concrete. For this reason, the bond stress distribution for bigcircular specimens is now quite different from that shown inFig. 5(a). Owing to the existence of the gap, there was no bond

stress formed near the loaded end initially, and the longitudinaldistribution of bond stress became much more uniform when τ u

was reached compared with that for specimen C9-90b shown inFig. 5(a).

It is worth noting that there is an exception to this size effectobserved for specimens C4-90a and C4-90b, which had evenhigher bond strength (increased by 17.9%) compared with theircounterparts C1-90a and C1-90b. A possible explanation for thisis that macro-interlocking effect was more significant for C4-90aand C4-90b since they had a much longer length than C1-90a andC1-90b. This phenomenon was not observed for corresponding

square columns since they hadmuch smaller effective contact areaand thus the influence of macro-interlocking was less significant.




Table 5

Design bond strength (τ Rd) specified in different design codes.

Code AIJ [21] AS 5100 [22] EC4 [23] BS 5400 [24] AISC [25] DBJ/T13-51-2010 [26]

Circular CFST 0.225 0.4 0.55 0.4 0.4 0.225

Rectangular CFST 0.15 0.4 0.4 0.4 0.4 0.15


Fig. 17. Effect of Li/D or Li/B ratio on bond strength.

Furthermore, compared with the reference specimens C4-0a andC4-0b, C4-90a and C4-90b also had higher bond strength. Thisis attributed to the fact that the gap between the steel tube andconcrete resulted from concrete shrinkage in large-size columnswas offset by the residual expansion of concrete in C4-90a and C4-90b. Therefore, bond stress due to friction developed very fast forthese specimens. But this is not the case for specimens C4-0a andC4-0b without fire exposure.

Currently, different design bond strengths (τ Rd) for CFSTcolumns are specified in different design codes as summarised inTable 5, i.e. AIJ [21], AS 5100 [22],EC4[23],BS 5400 [24], AISC[25],DBJ/T13-51-2010 [26]. Within the limitations of the current tests,0.4 MPa and 0.15 MPa can be adopted asτ Rd for circular and squareCFST columns, respectively. From the comparison shown in Fig. 15,there is no need to consider the fire exposure influence if τ Rd

is so specified. It is worth noting that three codes including AS5100, AISC and BS 5400 have specified a same value of 0.4 MPafor both circular and square columns. Since the bond mechanismsfor circular and square columns are quite different, the assignedvalue of 0.4 MPa for square columns often exceeds measured bondstrength, especially for those big columns. Actually, in this paper,0.4 MPa is the upper bound for both the unheated specimens andthose with a 90 min fire exposure as shown in Fig. 15.

3.5.3. Effect of Li/D or Li/B ratio

The relationship between τ u and Li/D or Li/B ratio is shownin Fig. 17. For circular columns, the trend is that τ u decreased

with increasing Li/D ratio. But the trend for square columns is just the opposite. Since bond strength is mainly provided by thesteel–concrete interaction near the loaded end for circularcolumns, the average bond strength will decrease if the interfacelength increases. Compared with circular tubes, the bond stressesdistribute more evenly along the length of square tubes. Theincrease of average bond strength as a function of Li/B ratio forsquare columns could be explained by the increasing influenceof macro-interlocking. The longer the column, the higher thetolerance of surface irregularities and overall imperfections are.

3.5.4. Effect of concrete type

Two types of concrete, i.e., SCC and NC, were used in the currenttest program. The effect of concrete type on τ u can be found in

Table 1. It seems that SCC can have a higher or lower bond strengththan NC, which greatly depends on the type of FA, water-binder

ratio, and cement replacement ratio used for the SCC. In general,the bond strength of NC was bigger than that of SCC when Class IIIFA was used, but was smaller than the bond strength of SCC whenClass II FA was used. Therefore, within the limitation of the currenttests, the bond strength of SCC is comparable to that of NC.

3.5.5. Effects of fly ash type and different SCC mixtures

Generally, Class II FA concrete had higher bond strength thanClass III FA concrete except specimens C2-0 and C2-90. Since ClassII FA is finer and more active than Class III FA, concrete madefrom Class II FA is expected to have higher chemical cohesion andresidual strength, which can enhance the bond behaviour betweenthe steel tube and concrete. For series C2, the concrete used had

higher strength and smaller water-binder ratio when comparedwith the concrete used in series C3. Therefore, the bond strengthof C2-0 and C2-90 was greatly affected by concrete shrinkage,and the gap thus had an adverse effect on the friction. Due tothe residual expansion of concrete, there was no gap influence forspecimen C2-180. Forthe square columns, their bond strength wasless sensitiveto concrete shrinkage. Hence,no exceptionwas foundfor the square columns with respect to the influence of FA type.

It can also be found from Table 1 that cement replacement ratioand water-binder ratio have some influence on τ u . For the circularspecimens,bond strength increased when the cementreplacementratio increased from 31% to 48% and water-binder ratio increasedfrom 0.31 to 0.38. But the trend was just the opposite for thesquare specimens. Undoubtedly, the concrete strength decreased

with increasing cement replacement ratio or water-binder ratio.This might also result in the reduction in chemical adhesion whichcontributed greatly to thebond strength of thesquare columns. Butfor the circular columns, higher concrete shrinkage resulted fromthe decrease of cement replacement ratio and water-binder ratioled to a decreased bond strength.

3.5.6. Effect of concrete curing condition

Concrete moisture could be prevented from dissipating at theends if a tube was sealed. Generally, smaller shrinkage can be ex-pected for concrete under sealed condition and reduce the pos-sible gap formed between the steel tube and concrete. It can befound from Table 1 that all the four sealed specimens with 180

min fire exposure had higher bond strength than their referencespecimens. The increases of average bond strength for the circular




and square columns were 13% and 49%, respectively. But this phe-nomenon was not observed for the two sealed square specimenswith 90 min fire exposure, and only one circular specimen C7-90bshowed extremely high bond strength which was different fromits companion specimen C7-90a. Therefore, a total of three sealedspecimens with a fire exposure of 90 min had lower τ u comparedwith the unsealed ones. At this moment, the variation in the testobservations is hardto interpret.More experimental investigationsare needed to clearly understand the mechanisms.

4. Conclusions

An experimental program of bond characteristics between thesteel tube and in-filled concrete is discussed in this paper. The fol-lowing conclusions can be drawn based on the results of this study.

(1) Bond strength generally decreased for specimens after 90 minfire exposure, and a strength recovery was found when the fireexposure time was extended to 180 min.

(2) Circular columns generally had much higher bond strengththan square columns. Within the limitations of the currenttests, 0.4 and 0.15 MPa are recommended as design bond

strength for circular and square CFST columns, respectively. Inthis regard, the influence of fire exposure may be ignored.

(3) Due to concrete shrinkage, bond strength of CFST columnswas very sensitive to their cross-sectional dimension. Moreresearch is needed for columns with large cross-sections, andmaybe suitable measures should be taken to improve the bondstrength.

(4) Fly ash type, water-binder ratio and cement replacement ratiohave impact on the bond strength of SCC. In general, the bondstrength of SCC is comparable to that of normal concrete.

Based on the research results reported in this paper, it seems thatfire exposure has significant influence on the initial slope, ultimatestrength, and curve shape of the bond stress-slip curves for CFSTcolumns. To carry out accurate nonlinear analysis for fire-damaged

structures, further research is needed to put forward a suitablebondstress-slip model incorporating the influence of fire exposure.The experimental work presented in this paper has provided agood basis for the development of the model.

Acknowledgements

The research reported in the paper is part of the NewCentury Outstanding Talents Supporting Project of Universities inFujian Province (XSJRC2007-05) and the Research Fund for theDoctoral Program of Higher Education of China (200803860005).This research work has also been partially supported by theAustralian Research Council (ARC) under its Future Fellowshipsscheme(Project No: FT0991433). The financial support is gratefully

acknowledged.

References

[1] Tao Z, Uy B, Han LH, Wang ZB. Analysis and design of concrete-filled stiffenedthin-walled steel tubular columns under axial compression. Thin-WalledStructures 2009;47(12):1544–56.

[2] Wang YC.Steel andcomposite structures: behaviourand designfor fire safety.London: Spon Press; 2002.

[3] Han LH, Huo JS. Concrete-filled hollow structuralsteel columns afterexposureto ISO-834 fire standard. Journal of Structural Engineering, ASCE 2003;129(1):

68–78.[4] ChiangCH, Tsai CL.Time-temperatureanalysisof bond strengthof a rebarafterfire exposure. Cement and Concrete Research 2003;33(10):1651–4.

[5] Hunaiti YM. Bond strength in battened composite columns. Journal of Structural Engineering, ASCE 1991;117(3):699–714.

[6] Virdi KS, Dowling PJ. Bond strength in concrete filled steel tubes. IABSEProceedings 1980;P-33/80:125–39.

[7] Shakir-Khalil H. Resistanceof concrete-filled steel tubesto pushout forces. TheStructural Engineer 1993;71(13):234–43.

[8] Roeder CW, Cameron B, Brown CB. Composite action in concrete filled tubes. Journal of Structural Engineering, ASCE 1999;125(5):477–84.

[9] Parsley MA, Yura JA, Jirsa JO. Push-out behavior of rectangular concrete-filledsteel tubes. In: Composite and hybrid systems. ACI Special Publication (SP-196), American Concrete Institute; 2000. p. 87–107.

[10] Nezamian A, Al-Mahaidi R, Grundy P. Bond strength of concrete plugsembedded in tubular steel piles under cyclicloading. Canadian Journal of CivilEngineering 2006;33(2):111–25.

[11] Nardin S, El Debs ALHC. Shear transfer mechanisms in composite columns: an

experimental study. Steel and Composite Structures 2007;7(5):377–90.[12] Aly T, Elchalakani M, Thayalan P, Patnaikuni I. Incremental collapse thresholdfor pushout resistanceof circular concrete filled steel tubular columns. Journalof Constructional Steel Research 2010;66(1):11–8.

[13] ISO 834. Fire-resistancetests—elements of building construction—part 1: gen-eral requirements, ISO 834-1. International organization for standardization.Geneva (Switzerland); 1999.

[14] GB/T 1596–2005. Fly ash used for cement and concrete; 2005 [in Chinese].[15] Tao Z, Han LH, Wang ZB. Experimental behaviour of stiffened concrete-

filled thin-walled hollow steel structural (HSS) stub columns. Journal of Constructional Steel Research 2005;61(7):962–83.

[16] Xue LH, Cai SH. Bond strength at the interface of concrete-filled steel tubularcolumns: part I. Building Science 1996;12(3):22–8 [in Chinese].

[17] Khoury GA. Strain of heated concrete during two thermal cycles. Part 1: strainover two cycles, during first heating and at subsequent constant temperature.Magazine of Concrete Research 2006;58(6):367–85.

[18] Schneider U. Concrete at high temperature—a general review. Fire Safety Journal 1988;13(1):55–68.

[19] ABAQUS. ABAQUSStandarduser’s manual, version 6.7.2.Providence (RI, USA):Dassault Systèmes Corp.; 2007.

[20] Han LH,Wang WH,Yu HX.Experimentalbehaviourof reinforcedconcrete (RC)beam to concrete-filled steel tubular (CFST) column frames subjected to ISO-834 standard fire. Engineering Structures 2010;32(10):3130–44.

[21] Architectural Institute of Japan (AIJ). Recommendations for design andconstruction of concrete filled steel tubular structures. Japan; 1997 [in Japanese].

[22] StandardsAustralia.AS5100.6-2004 Bridge design, part6: steel and compositeconstruction. Sydney (Australia); 2004.

[23] Eurocode 4. Design of compositesteel andconcretestructures,part1.1,generalrules and rules for building. BS EN 1994-1-1: 2004. London: British StandardsInstitution; 2004.

[24] BS5400. Steel, concrete and composite bridges, part 5, code of practice for thedesign of composite bridges. London (UK); 2005.

[25] ANSI/AISC 360-10. Specification for structural steel buildings. Chicago (IL,USA): American Institute of Steel Construction; 2010.

[26] DBJ/T13-51-2010. Technical specification for concrete-filled steel tubularstructures. Fuzhou (China): The Department of Housing and Urban-RuralDevelopment of Fujian Province; 2010 [in Chinese].

Documents

Post-Fire Bond Between the Steel Tube and Concrete in Concrete-filled Steel Tubular Columns