Experimental Test on SSTT-Confined HSC Columns

Magazine of Concrete Research

http://dx.doi.org/10.1680/macr.14.00065

Paper 1400065

Received 28/02/2014; revised 01/05/2014; accepted 12/05/2014

ICE Publishing: All rights reserved

Magazine of Concrete Research

Experimental tests on SSTT-confined HSCcolumnsMa, Awang, Omar and Maybelle

Experimental tests onSSTT-confined HSC columnsChau-Khun MaPhD Candidate, Faculty of Civil Engineering, Universiti Teknologi Malaysia,Johor Bahru, Malaysia

Abdullah Zawawi AwangSenior Lecturer, Faculty of Civil Engineering, Universiti Teknologi Malaysia,Johor Bahru, Malaysia

Wahid OmarProfessor, Faculty of Civil Engineering, Universiti Teknologi Malaysia, JohorBahru, Malaysia

Liang MaybellePhD Candidate, Faculty of Civil Engineering, Universiti Teknologi Malaysia,Johor Bahru, Malaysia

The steel-strapping tensioning technique (SSTT) has been widely accepted as an effective method for enhancing the

performance of high-strength concrete (HSC) columns. Previous experimental tests showed that SSTT can increase the

ductility of HSC by up to twice that of unconfined HSC. However, most of the tests performed on SSTT-confined HSC

columns have focused on concentrically loaded short specimens. In reality, however, columns with a length/diameter

ratio greater than three and subjected to eccentric loading are very common. Against this background, experiments

were carried out to investigate the slenderness effect of SSTT-confined HSC columns subjected to eccentric loads. It

was found that SSTT increases both the strength and deformability of slender HSC columns, although the confining

effects are reduced proportionally with an increase in slenderness ratio. The effects of the eccentricities and the

eccentricity ratio are also presented in this paper.

NotationAs column cross-sectional area

e eccentricity

ei /es eccentricity ratio

fc column strength

fy yield strength of steel straps

L total length of the column

r gross radius of gyration of the column cross-section

Vc volume of concrete being confined

Vs volume of steel straps used

�MID lateral mid-height deflection

º slenderness ratio (¼ L/r)

r longitudinal reinforcement ratio

rv SSTT confinement ratio

IntroductionHigh-strength concrete (HSC) has become popular in recent years

due to its high strength and dense microstructure. However, the

application of HSC in the construction industry is still rather

stagnant. Over the past 20 years, the construction of multi-storey

buildings in Malaysia has been very limited. The construction of

Petronas Twin Towers can be considered the most significant

breakthrough of the use of HSC in Malaysia, which used concrete

of grade 90 MPa for the basement storey. The limited use of

higher grade concrete in Malaysia, despite the rapid development

of concrete that can be commercially produced up to 300 MPa

strength, provides evidence of a lack of confidence in the use of

HSC. In view of the many positive findings elsewhere, the main

barrier to wider application of HSC in Malaysia may be ascribed

to the brittleness of the material: HSC, although offering higher

strength and durability, is very brittle compared with normal-

strength concrete (NSC). Indeed, it has been reported that the

ductility of HSC structures reduces almost linearly with an

increase in concrete strength and the failure of HSC can be

sudden and explosive.

A recent research study conducted at Universiti Teknologi

Malaysia on confined HSC showed that the use of low-cost steel

straps as confining material can significantly increase the strength

and ductility of HSC (Awang, 2013). The provision, by the steel

straps, of a lateral pre-tensioned force was found to be very

effective in restricting the lateral dilation of HSC and conse-

quently improved the cracking strain of extreme concrete fibres in

loaded HSC. The stress–strain curve of HSC confined with steel

straps showed strain softening and the inherently low deform-

ability of HSC was observed to increase with confinement ratio.

However, research studies on HSC columns confined with steel

straps have mainly focused on concentrically loaded short speci-

mens. Although some research studies have been carried out to

study the behaviour of eccentrically loaded columns confined

with steel straps (Hadi, 2011; Lei, 2012; Song, 2012), the effects

of slenderness and eccentricity of confined HSC are still poorly

understood.

In this work, tests were performed to

j evaluate the effectiveness of the steel-strapping tensioning

technique (SSTT) for medium-scale circular HSC columns of

increasing slenderness by comparison with identical

unconfined columns

1

j study the effectiveness of hoop-style SSTT-confinement in

reducing the susceptibility of slender HSC columns to

slenderness effects.

Experimental method

Details of column specimens

A total of 18 circular columns were prepared for testing. All the

columns were 150 mm in diameter and were 600, 900 or

1200 mm in length. The cross-sectional dimensions and reinforce-

ment details of the columns are shown in Figure 1. The long-

itudinal reinforcement consisted of four 10 mm diameter

deformed bars that were equally distributed around the column.

The selected target ratios of longitudinal reinforcement were

0.7% for the columns. The lateral reinforcement of the columns

consisted of 6 mm diameter deformed bars with a spacing of

300 mm. The internal reinforcements were chosen for construc-

tive purposes only according to EN 1992-1-1 (CEN, 1992). The

column dimensions and reinforcement details were limited to the

capacity and vertical dimensions of the loading frame used, and

also by the minimum sizes of available deformed steel bars.

Although the columns were slightly smaller than realistic

columns, studies have found no discernible size effect for circular

fibre-reinforced polymer (FRP) wrapped reinforced concrete

columns (Rocca et al., 2006).

The column details are summarised in Table 1. The clear

spacings of the steel straps were selected to be 20 mm and

40 mm, with one or two steel strap layers. The measured concrete

compressive strength, based on standard 100 mm 3 100 mm 3

100 mm cube tests, was in the range 60–70 MPa. Confinement

was provided up to the ends of the columns to avoid premature

failure in the end regions of the columns. The concrete cover was

kept constant at 20 mm to the face of the longitudinal reinforce-

ment. The longitudinal reinforcements were cut to match the

column height and were flush with the ends of the columns. The

circular columns were prepared for vertical casting using poly-

vinyl chloride (PVC) tubes. The PVC tubes were secured by

fitting them in plywood formworks (internal dimensions of

152 mm 3 152 mm) to avoid buckling of the tubes during

casting.

Specimen identification

Column specimens were labelled according to slenderness ratio

(i.e. different lengths of 600, 900 and 1200 mm), eccentricity of

applied load, eccentricity ratio and SSTT confinement ratio. For

example, specimen C600-E25-R1-S0.076 is a circular column of

600 mm length, subjected to eccentricities of 25 mm, with an

eccentric ratio of 1 and confined with a SSTT confinement ratio

(rv) of 0.076. The SSTT confinement ratio is calculated based on

Vsfy /VcAs, where Vs and Vc are the volume of steel straps used

and the volume of concrete being confined respectively, fy is the

yield strength of the steel straps and As is the column’s cross-

sectional area.

Steel-strapping method

The technique for strengthening the HSC column involves the

pre-tensioning of steel straps around the column (using standard

strapping machines used in the packaging industry) and securing

them in place with steel clips. Using commercially available

strapping equipment, the straps are tensioned around the column

by a pressure of approximately 0.2 MPa, as recommended by

Awang (2013), so that an effective lateral stress can be applied on

the column prior to loading. This ensures effective utilisation of

the straps and avoids early crushing of the confined concrete,

which can occur without properly tightened straps. Table 2 gives

the specifications of the PT-52 pneumatic tensioner and the steel

straps and the pneumatic tensioner are shown in Figure 2. The

low cost of the straps and the speed and ease of application of

the strapping technique make this an efficient repair and

strengthening technique for certain structural members (Moghad-

dam et al., 2008).

Test set-up

The column specimens were hinged at one end with the other end

placed on flat ground. Load was applied with an initial eccen-

tricity of 25 mm at one end. The hinge was formed by a steel

knife-edged bearing plate in contact with a 40 mm thick steel

plate fixed to the column end. All columns were instrumented

with strain gauges (6 mm) in the hoop direction at mid-height.

Lateral deflection was monitored using three linear variable

differential transformers (LVDTs), one at the mid-height of the

specimens and the others at 1/4 of the total length of the

A ASection A–A

110 mm150 mm

R6 – 300 mm/C/C

Concrete 60 MPa

0·5 mm thick/15 mmwide steel straps

10 mm diameterlogitudinal bars

Figure 1. Detail of test specimens

2

Magazine of Concrete Research Experimental tests on SSTT-confined HSCcolumnsMa, Awang, Omar and Maybelle

specimens from both ends (Figure 3). A Dartec 2000 kN com-

pression testing machine was used in all loading tests.

All column specimens were cleaned and ground to ensure an

even distribution of forces. The specimen to be tested was then

lifted up vertically and placed into the eccentricity cap. The same

procedure was applied to the other end of the column after curing

for about 30 min. After both ends were capped and fixed, the

column was mounted on the testing machine. Calibration was

carried out to ensure the column specimen was centrally placed

by checking the centreline using a levelling rule. Once the

column specimen was centred, the actuator head was lowered

slowly to contact with the eccentric knife plate. This step was

important to ensure that both sides of the eccentric knife plate

were cleared from touching the eccentric head as this would lead

to different eccentricity conditions. Testing was then started with

an applied load rate of 4 mm/s until failure.

Test results

Unconfined columns

As expected, the unconfined columns exhibited sudden failure. The

failure mode of the confined columns was explosive near the

ultimate load. Figure 4 shows typical failure modes for unconfined

columns. Failure of the columns was mainly triggered by crushing

of the concrete in the middle of the columns at the most

compressed side. For column C600-E25-R0-U, failure appeared to

coincide with localised buckling of compression longitudinal steel

bars due to spalling of the concrete cover at approximately mid-

height. For column C900-E25-R0-U, failure occurred when cracks

initiated from the top and mid-height of the column coincided

(Figure 4(b)): failure of this column was very sudden with a loud

crushing sound. Column C1200-E25-R0-U showed better deform-

ability than the other two specimens, but failure was also accom-

panied by immediate concrete crushing and a loud crushing sound.

In general, the failures of all the unconfined columns were sudden

and explosive, with the result that the post-peak behaviour of the

unconfined columns was not recorded in the load–deflection

relationships. In Figure 5, only the lateral mid-height deflection

�MID is considered and is termed deflection here for brevity. The

Model PT-52

Type Pneumatic pressure

Maximum tension: kg 900

Maximum pressure: MPa 0.6

Suitable strap thickness: mm 0.6–1.2

Maximum strap width: mm 32

Table 2. Tensioner specifications

Column ID L: mm fc: MPa Longitudinal reinforcement rv e: mm ei/es Number of layers

Number and

size: mm

r: % fy: MPa

C600-E25-R0-U 600 61.5 4 ˘10 0.02 460 — 25 0 Unconfined

C600-E25-R0-S0.12 600 61.5 4 ˘10 0.02 460 0.120 25 0 1

C600-E25-R0-S0.076 600 61.5 4 ˘10 0.02 460 0.076 25 0 1

C600-E25-R0-S0.178 600 61.5 4 ˘10 0.02 460 0.178 25 0 2

C600-E50-R0-S0.012 600 61.5 4 ˘10 0.02 460 0.120 50 0 1

C600-E25-R1-S0.012 600 61.5 4 ˘10 0.02 460 0.120 25 1 1

C900-E25-R0-U 900 67.3 4 ˘10 0.02 460 — 25 0 Unconfined

C900-E25-R0-S0.12 900 67.3 4 ˘10 0.02 460 0.120 25 0 1

C900-E25-R0-S0.076 900 67.3 4 ˘10 0.02 460 0.076 25 0 1

C900-E25-R0-S0.178 900 67.3 4 ˘10 0.02 460 0.178 25 0 2

C900-E50-R0-S0.12 900 67.3 4 ˘10 0.02 460 0.120 50 0 1

C900-E25-R1-S0.12 900 67.3 4 ˘10 0.02 460 0.120 25 1 1

C1200-E25-R0-U 1200 63.4 4 ˘10 0.02 460 — 25 0 Unconfined

C1200-E25-R0-S0.12 1200 63.4 4 ˘10 0.02 460 0.120 25 0 1

C1200-E25-R0-S0.076 1200 63.4 4 ˘10 0.02 460 0.076 25 0 1

C1200-E25-R0-S0.178 1200 63.4 4 ˘10 0.02 460 0.178 25 0 2

C1200-E50-R0-S0.12 1200 63.4 4 ˘10 0.02 460 0.120 50 0 1

C1200-E25-R1-S0.12 1200 63.4 4 ˘10 0.02 460 0.120 25 1 1

Table 1. Specimen details

3


load–deflection relationships of unconfined columns with various

slenderness ratios (º ¼ 16, 24 and 32) are shown in Figure 5. The

slenderness ratio is taken as L/r, where L is the total length of the

column and r is the gross radius of gyration of the column cross-

section. It was found that an increase in slenderness resulted in

decreased ultimate load-carrying capacity but an increase in

deflection. This indicates that the failure mode shifted from

axially dominated material failure to flexural dominated instabil-

ity failure. Figure 6 suggests that no yielding of the longitudinal

bars occurred prior to failure and therefore the ultimate deflec-

tions of all the unconfined columns were relatively small.

Examination of the relationships of load against concrete axial

strain (Figure 7) reveals that both tensile concrete strain and

compressive concrete strains presented almost similar patterns for

columns with various slenderness ratios when subjected to

eccentric loads. The variation of axial strain with increasing

slenderness for unconfined HSC columns is counterintuitive as no

significant increments were observed for axial concrete strains

with increasing flexural effects. This can be explained by noting

that failure of the unconfined slender HSC columns was initiated

slightly above the mid-height of the specimens, where the strain

gauges were placed. This led to inaccuracy as the maximum

strain was not captured at the exact location of failure in most

cases (which is at mid-height of the column).

Confined columns

Figure 8 shows typical failure modes of the eccentrically loaded

SSTT-confined slender HSC columns. It can be seen that failure

Load actuator

Load cell

LVDT 50 mm

LVDT 50 mm

LVDT 50 mm

L /4

L /4

Pi gauges

Strain gauges

Figure 3. Schematic illustration of test set-up and selected

instrumentation (A ‘pi gauge’, also known as an omega strain

gauge, is a non-destructive gauge used to measure a material’s

strain. Its shape resembles the Greek character pi/omega – hence

its name.)

(a)

(b)

Figure 2. Steel straps (a) and tensioner (b) used to confine the

columns

4


of the columns changed from compression-sided failure to

tensile-sided failure after confinement. Failures were noted to

have occurred at the tension side, at approximately mid-height.

Distinct horizontal cracks were observed at the mid-height of

some of the columns. Large deformations were observed at

ultimate failure and most of the confined columns failed in a

more gentle manner than the unconfined columns. Some of the

steel straps near the column ends had snapped at ultimate load,

but none of the steel straps around the mid-height of the columns

were snapped, as can be seen in Figure 8. This is because the

different end eccentricities applied at the ends of the specimens

led to unsymmetrical bending of specimens.

The load–deflection relationships shown in Figure 9 suggest that

the SSTT-confined HSC columns displayed an initial linear elastic

stage followed by a gradual reduction in flexural stiffness and a

subsequent essentially linear ascending or descending branch

until failure depending on the column’s slenderness. Failures of

the confined columns were observed either at the peak axial load

for cases with no descending branch or at a reduced post-peak

load with large deformation. As expected, the confined columns

exhibited higher axial and deformability than their unconfined

counterparts, able to sustain comparatively large lateral deflec-

tions at ultimate load. However, increments in axial load and

deformability varied according to the slenderness ratio and SSTT

confinement ratio. Increased slenderness resulted in decreased

ultimate axial capacity and increased lateral deflection at ultimate

as the behaviour became more flexure-dominated due to the

amplification of secondary moments.

Figure 10 shows selected axial load against longitudinal steel

strain curves for SSTT-confined HSC columns (eccentricity

e ¼ 25 mm, eccentricity ratio ei /es ¼ 0, SSTT confinement ratio

rv ¼ 0.178). The figure shows that the compressive steel yielded

at ultimate load only for the column with º ¼ 16. This is because

the stresses were concentrated on the compression side during

application of loading. However, with increasing slenderness, the

failure modes tended to shift from the compression side to the

tension side as smaller steel strains were observed for the

columns with º ¼ 24 and 32. Interestingly, for all the confined

columns, the tensile longitudinal steel strains were compressed up

to approximately 80% of ultimate load. The tensile steels were

then tensioned after this point. This is completely different from

what was observed in the unconfined columns as their long-

itudinal tensile steels were tensioned from the start of the load

application. This may be due to SSTT-confinement effects and

more detailed investigations are needed in this area.

(a) (b) (c)

Figure 4. Typical failure modes of unconfined columns:

(a) C600-E25-R0-U, (b) C900-E25-R0-U and (c) C1200-E25-R0-U

2·52·01·51·00·50

100

200

300

400

500

600

700

800

0

Axi

al lo

ad,

: kN

N

Mid-height deflection, : mmδMID

C600-E25-R0-UC900-E25-R0-UC1200-E25-R0-U

λ � 32

λ � 24

λ 16�

Figure 5. Load–deflection relationships for unconfined columns

(e ¼ 25 mm, ei/es ¼ 0, rv ¼ 0)

5


Figure 11 shows selected load against axial concrete strain curves

for SSTT-confined HSC columns (e ¼ 25 mm, ei /es ¼ 0,

rv ¼ 0.178). Again, there are secondary ascending or descending

branches depending on the slenderness of the specimen. The

axial concrete strains throughout the cross-sections of the

confined columns were considerably higher than those of the

unconfined specimens. This suggests that SSTT confinement can

improve the axial concrete strain of HSC columns, and this was

also discussed in other work on concentrically loaded FRP-

confined concrete (Bisby et al., 2005). Bisby et al. (2005)

concluded that the effectiveness of FRP confinement was drasti-

cally reduced with an increase in slenderness. A similar observa-

tion can be found in Figure 11 as axial concrete strains

diminished with increased º.

Analysis and discussion

Effect of slenderness

Figure 12 shows the effect of increasing slenderness on the lateral

mid-height deflection on both unconfined and confined HSC

columns and Figures 13 and 14 show the effects of increasing

slenderness on the ultimate strength and axial concrete strain

respectively (linear least-squares trend lines are shown in Figure

13 for illustrative purposes). As expected, an increase in slender-

ness caused the ultimate load capacity to decrease and increased

Axi

al lo

ad,

: kN

N

Yield strain 0·0021�

0·00250·00150·0005�0·0005�0·00150

100

200

300

400

500

600

700

800

�0·0025Steel strain: mm/mm

C600-E25-R0-U

C900-E25-R0-U

C1200-E25-R0-U

Yield strain 0·0021� �

Figure 6. Load–longitudinal steel strain relationships for

unconfined columns (negative strain values indicate tension)

0·00100·00080·00060·00040·00020�0·0002�0·0004�0·00060

100

200

300

400

500

600

700

800

�0·0008

Axi

al lo

ad,

: kN

N

Concrete axial strain, : mm/mmεc

C600-E25-R0-U

C900-E25-R0-U

C1200-E25-R0-U

Figure 7. Load–concrete axial strain relationships (negative strain

values indicate tension)

6


the lateral mid-height deflection of the confined columns. Deflec-

tion can be considered as one of the indications of deformability

or displacement ductility. An increase in º clearly led to increased

deflection, but there was no distinct difference between the

increases in HSC columns confined with SSTT with confinement

ratios rv of 0 to 0.178 (rv ¼ 0 represents the unconfined column):

for an increase in slenderness ratio from 16 to 32, the increase in

deflection was �237% for the unconfined column and �250%

for the confined column with rv ¼ 0.178.

Slenderness effects on ultimate load capacity were more signifi-

cant for higher SSTT confinement ratios, as can be seen in Figure

13. The columns were flexure-dominated for a higher slenderness

ratio and lateral hoop confinement pressure was ineffective in

significantly increasing the flexural rigidity of slender columns

(Jiang and Teng, 2012a, 2012b; Tamuzs et al., 2007). As the

slenderness ratio increased, loss of ultimate load capacity seems

to accelerate with an increase in SSTT confinement ratio. A

similar observation was reported for slender FRP-confined col-

umn tests (Mirmiran et al., 2001; Yuan and Mirmiran, 2001;

Yuan et al., 2008). SSTT confinement increased the serviced load

significantly without effectively improving the flexural rigidity of

the columns.

The ineffectiveness of SSTT confinement in increasing the

flexural rigidity of HSC columns is illustrated in Figure 14. This

(a) (b) (c)

(d) (e) (f)

Figure 8. Typical failure modes of eccentrically loaded confined

columns

7


figure shows the effect of slenderness on the axial load against

ultimate axial concrete strain curves for both unconfined and

confined specimens. For columns with º ¼ 16, the ultimate axial

concrete strain is greater for higher values of rv. However, for all

cases (unconfined and confined columns), the ultimate axial

concrete strain reduces drastically between º ¼ 16 and º ¼ 24.

This is due to the ineffectiveness of SSTT at improving the axial

concrete strain for flexure-denominated/slender columns. For the

4·03·53·02·52·01·51·00·50

200

400

600

800

1000

1200

0

Axi

al lo

ad,

: kN

N

Mid-height deflection, : mmδMID

λ 16�

λ � 24

λ � 32

Figure 9. Load–deflection curves for confined columns

(e ¼ 25 mm, ei/es ¼ 0, rv ¼ 0.178)

0·00300·00250·00200·00150·00100·00050

Axi

al lo

ad,

: kN

N

Yield strain0·0021�

0100200300400500600700800900

1000

�0·0005Steel strain: mm/mm

λ 16�

λ � 24

λ � 32

Figure 10. Axial load–longitudinal steel strain curves for confined

columns (negative strain values indicate tension)

0·00250·00200·00150·00100·000500

100200300400500600700800900

1000

�0·0005

Axi

al lo

ad,

: kN

N

Concrete strain: mm/mm

λ � 16λ � 24

λ 32�

Figure 11. Axial load–concrete axial strain curves (negative strain

values indicate tension)

Mid

-hei

ght

defle

ctio

n,: m

mδ M

ID

353025200

1

2

3

4

15Slenderness ratio, λ

Unconfined

ρv 0·076�

ρv � 0·120

ρv � 0·178

Figure 12. Effect of slenderness ratio and SSTT confinement ratio

on lateral mid-height deflection (specimen group with e ¼ 25 mm

and ei/es ¼ 0)

333129272523211917500

600

700

800

900

1000

1100

1200

15

Ulti

mat

e lo

ad c

apac

ity: k

N

Slenderness ratio, λ

Unconfined

ρv � 0·076

ρv 0·120�

ρv � 0·178


on ultimate load capacity (specimen group with e ¼ 25 mm and

ei/es ¼ 0)

353025200

0·001

0·002

0·003

0·004

0·005

0·006

15

Con

cret

e st

rain

: mm

/mm

Slenderness ratio, λ

Unconfined

ρv 0·076�

ρv � 0·120

ρv � 0·178


on axial compressive strain (specimen group with e ¼ 25 mm and

ei/es ¼ 0)

8


case of rv ¼ 0.178, there is an approximately 77% drop in

ultimate axial concrete strain for º increasing from 16 to 24. The

ultimate concrete axial strains for all confined and unconfined

specimens beyond º ¼ 24 were very similar. This clearly explains

the flexure-dominated behaviour and under-utilisation of SSTT

confinement in slender columns.

Effect of SSTT confinement ratio

Figure 15 shows the effect of increasing SSTT confinement ratio

(rv ¼ 0, 0.076, 0.120, 0.178) on the load–deflection curves

behaviour of both short and slender columns (rv ¼ 0 represents

an unconfined column). Increasing rv increased the strength and

deformability for both short and slender columns under eccentric

axial load, but the increase in axial load is more pronounced for

the short columns. The increments in axial load were approxi-

mately 30% and 80% for columns with rv ¼ 0.076 and 0.120

respectively. It is probable that the pre-tensioned force provided

by SSTT confinement improves the concrete’s strength and

deformability mainly by preventing or delaying local bar buckling

for a higher SSTT confinement level. The enhanced axial con-

crete strain, although not significant, may also play a role.

Based on Figure 16, the effects of SSTT confinement ratio are

most pronounced for columns with º ¼ 16. Columns with º ¼ 16

showed an approximately 71% increase in mid-height deflection

compared with the unconfined specimen, but the short columns

(º ¼ 16) showed an increase of only 52% compared with

unconfined specimen. Although the improvement in deformability

is less for short columns, a minimum of 50% improvement is

guaranteed. However, in terms of improvement in axial concrete

compressive strain, much better performance was observed for

the short columns. As shown in Figure 17, compared with

unconfined columns, an increase in ultimate axial concrete strain

of approximately 689% was observed for º ¼ 16, whereas the

increase for columns with º ¼ 24 and 32 was just 83%.

Obviously, SSTT confinement is less effective in confining more

slender columns. This again confirms that hoop-style SSTT

confinement is not that effective in increasing the flexural rigidity

of slender HSC columns.

Effect of load eccentricity

Figure 18 shows the influence of eccentricity on the ultimate load

of SSTT-confined HSC columns with slenderness ratios of 16, 24

and 32. The figure shows only the specimens confined with

rv ¼ 0.120 (confined with one layer of steel straps with a clear

spacing of 20 mm) but subjected to eccentricities of 25 mm and

50 mm. It is clear that column capacity is highly affected by the

level of load eccentricity. Although the same trend was observed

for all three groups (º ¼ 16, 24 and 32), as eccentricity increases,

the column capacities of columns with higher slenderness ratio

decrease more rapidly than the lower slenderness ratio columns.

The ultimate capacity of º ¼ 16 columns decreased by 34% as

the eccentricity increased from 25 mm to 50 mm. For columns

Axi

al lo

ad,

: kN

N

43210

200

400

600

800

1000

1200

0Mid-height deflection, : mmδMID

C600-E25-U

C600-E25-1L 40-

C600-E25 1L 20- -

C600-E25 2L 40- -

C1200-E25-U

C1200-E25 1L 40- -

C1200-E25 1L 20- -

C1200-E25 2L 40- -

Figure 15. Effect of SSTT confinement level on short and slender

specimens

0·200·150·100·050

1

2

3

4

0

Mid

-hei

ght

late

ral d

efle

ctio

n,: m

mδ M

ID

SSTT-confinement ratio, ρv

λ 16�

λ � 24λ � 32

Figure 16. Effect of SSTT confinement level on lateral mid-height

deflection

0·200·150·100·050

0·001

0·002

0·003

0·004

0·005

0·006

0

Axi

al c

oncr

ete

stra

in: m

m/m

m

SSTT-confinement ratio, ρv

λ 16�

λ � 24

λ � 32

Figure 17. Effect of SSTT confinement level on axial concrete

strain

9


with º ¼ 32, the reduction in ultimate capacities was as high as

60%.

Figure 19 shows that an increase in load eccentricity increases

the lateral mid-height deflection of the columns. The eccentricity

effect is insignificant for short columns, but an 82% increase in

deflection was recorded for columns with º ¼ 32. Large deflec-

tion can lead to earlier loss of ultimate capacity of columns due

to the second-order effect. This explains why the slender columns

(º ¼ 32) experienced a larger ultimate capacity loss than short

columns (º ¼ 16).

Effect of eccentricity ratio

Table 3 summarises the effects of eccentricity ratio on ultimate

load capacity, deflection and axial concrete strain. It should be

noted that the test results are based on all specimens confined

with a confinement ratio rv ¼ 0.120. It is clear that the ultimate

load capacity of SSTT-confined columns is affected by the

eccentricity ratio ei /es and the effect is more pronounced for

columns with a high slenderness ratio (º ¼ 16). A clear descend-

ing trend can also be observed in the lateral mid-height deflection

of SSTT-confined columns as slenderness ratio increased. This is

easily understood as a column bent in symmetrical curvature

(with ei/es ¼ 1) always experiences the largest slenderness effect.

The effect of eccentricity ratio on axial concrete strain is not that

significant compared to its effect on ultimate load capacity and

lateral mid-height deflection.

Conclusionsj The NSC columns fail in a more ductile manner with an

increase in slenderness ratio. However, HSC columns were

observed to fail in a brittle and sudden manner but with an

increased slenderness ratio.

j With use of the steel-strapping tensioning technique (SSTT),

the HSC columns failed in a more ductile manner.

j SSTT confinement can increase the strength and

deformability of both short and slender HSC columns, but the

effectiveness of SSTT confinement decreases with an increase

in slenderness ratio.

j The beneficial effects of SSTT confinement appear to be

greater for a higher confinement ratio. This may be

influenced by the observed axial strain enhancement or by the

pre-tensioned force provided by SSTT preventing local

buckling of the longitudinal reinforcing bars and spalling of

the concrete cover.

j SSTT confinement is less effective in increasing the ultimate

load capacity of slender HSC columns due to its

ineffectiveness in improving flexural rigidity/stiffness.

55504540353025Eccentricity, : mme

300

400

500

600

700

800

900

20

Ulti

mat

e lo

ad,

: kN

N

λ 16�

λ � 24

λ � 32

Figure 18. Effect of load eccentricity on ultimate load capacity

Eccentricity, : mme55504540353025

0

1

2

3

4

5

6

20

Def

lect

ion,

: mm

δ MID

λ 16�

λ � 24

λ � 32

Figure 19. Effect of load eccentricity on lateral mid-height

deflection

º Eccentricity

ratio, ei/es

Ultimate load,

N: kN

Deflection,

�MID: mm

Axial concrete strain,

�: mm/mm

16 0 860.0 0.80 0.0027

16 1 835.3 0.89 0.0022

24 0 825.4 1.48 0.0011

24 1 792.8 2.32 0.0010

32 0 810.0 2.87 0.0011

32 1 535.5 4.96 0.0008

Table 3. Effects of eccentricity ratio on ultimate load capacity,

deflection and axial concrete strain

10


REFERENCES

Awang AZ (2013) Stress–Strain Behavior of High-Strength

Concrete with Lateral Pre-Tensioning Confinement. PhD

dissertation, Universiti Teknologi Malaysia, Johor Bahru,

Malaysia.

Bisby LA, Dent AJS and Green MF (2005) Comparison of

confinement models for fiber-reinforced polymer-wrapped

concrete. ACI Structural Journal 102(1): 62–72.

CEN (European Committee for Standardization) (1992) ENV

1992-1-1. Eurocode 2: Design of concrete structures. Part 1:

General rules and rules for Building. European Committee

for Standardization, Brussels, Belgium.

Hadi MNS (2011) External reinforcement of concrete columns.

Proceedings of 11th Underground Coal Operators’

Conference, University of Wollangong, NSW, Australia,

Australian Institute of Mining and Metallurgy, Carlton,

Victoria, Australia, pp. 121–128.

Jiang T and Teng JG (2012a) Behavior and design of slender

FRP-confined circular RC columns. Journal of Composites

for Construction 16(6): 650–661.

Jiang T and Teng JG (2012b) Slenderness limit for short FRP-

confined circular RC columns. Journal of Composites for

Construction 17(4): 443–453.

Lei X (2012) Shape Modification of Square Reinforced Concrete

Columns Confined with Fibre-Reinforced Polymer and Steel

Straps. MEng thesis, University of Wollongong, Hong Kong.

Mirmiran A, Shahawy M and Beitleman T (2001) Slenderness

limit for hybrid FRP-concrete columns. Journal of

Composites for Construction 5(1): 26–34.

Moghaddam H, Pilakoutas K, Samadi M and Mohebbi S (2008)

Behaviour and modeling of concrete columns confined by

external post-tensioned strips. Proceedings of Structures

Congress 2008, Reston, Virginia, USA. ASCE, Reston, VA,

USA, pp. 1–10.

Park R and Paulay T (1975) Reinforced Concrete Structures.

Wiley, New York, NY, USA.

Rocca S, Galati N and Nanni A (2006) Large-size reinforced

concrete columns strengthened with carbon FRP:

experimental evaluation. Proceedings of 3rd International

Conference on FRP Composites in Civil Engineering (CICE

2006), Miami, FL, USA.

Song YM (2012) Behavior of Confined High-Strength Concrete

Column Subjected to Concentric and Eccentric Loads. MEng

thesis, Universiti Teknologi Malaysia, Johor Bahru, Malaysia.

Tamuzs V, Tepfers R, Zile E and Valdmanis V (2007) Stability of

round concrete columns confined by composite wrappings.

Mechanics of Composite Materials 43(5): 445–452.

Yuan W and Mirmiran A (2001) Buckling analysis of concrete-

filled FRP tubes. International Journal of Structural Stability

and Dynamics 1(3): 367–383.

Yuan XF, Xia SH, Lam L and Smith ST (2008) Analysis and

behaviour of FRP-confined short concrete columns subjected

to eccentric loading. Journal of Zhejiang University Science

A 9(1): 38–49.

WHAT DO YOU THINK?

To discuss this paper, please submit up to 500 words to

the editor at [email protected]. Your contribution will

be forwarded to the author(s) for a reply and, if

considered appropriate by the editorial panel, will be

published as a discussion in a future issue of the journal.

11


Documents

Experimental Test on SSTT-Confined HSC Columns