slag concrete members under combined loading

Engineering Structures 199 (2019) 109598

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Glass fibre-reinforced polymer circular alkali-activated fly ash/slag concrete members under combined loading Minhao Donga, Mohamed Elchalakanib, Ali Karrechc, Thong M. Phamd, Bo Yange,

T

⁎

a

School of Civil, Environmental and Mining Engineering, The University of Western Australia, WA 6009, Australia School of Civil, Environmental and Mining Engineering, The University of Western Australia, WA 6009, Australia c School of Civil, Environmental and Mining Engineering, The University of Western Australia, WA 6009, Australia d Center for Infrastructural Monitoring and Protection, School of Civil and Mechanical Engineering, Curtin University, Kent Street, Bentley, WA 6102, Australia e School of Civil Engineering, Chongqing University, Chongqing 400045, China b

A R T I C LE I N FO

A B S T R A C T

Keywords: Circular cross-section GFRP Spirals Alkali-activated fly ash/slag concrete Combined loading Interaction diagram

Glass fibre reinforced polymer (GFRP) reinforced alkali-activated fly ash/slag concrete (AAFS) could potentially be used as a durable construction material, especially in aggressive environments. This study aims to examine the structural behaviour of circular GFRP-reinforced AAFS members under combined loading, and introduces a geometrical factor for predicting the load and moment interaction of circular members. A total of 17 reinforced AAFS columns and beams (14 with longitudinal GFRP bars and 3 with steel bars) was constructed. GFRP spirals were used in all the specimens as transverse reinforcement. The effect of the number of longitudinal bars, the spiral pitch, the material of longitudinal bars and load eccentricities on confinement, load and moment capacity and failure modes was investigated. The GFRP reinforcement resulted in slightly reduced load capacities of 10%, however with an average improved ductility of 24% compared to steel reinforcement. The confinement was found to be affected by a combination of longitudinal and transverse reinforcement. The differences in load capacities between columns with 40 mm and 80 mm spiral pitches gradually diminished as the eccentricity increased. Interaction diagrams were constructed and validated based on the experimental data. The introduced geometrical factor could be used safely and effectively to obtain the equivalent stress block of circular sections. The calculation with a smaller stress block factor was the most accurate to the experimental results of AAFS members due to their reduced stress block size.

1. Introduction The use of alkali-activated material (AAM) and glass fibre reinforced polymer (GFRP) bars has been increasingly researched due to their advantageous properties over their conventional steel-reinforced ordinary Portland cement (OPC) concrete. The binder mainly consists of industrial by-products such as fly ash and ground granulated blastfurnace slag (GGBS), which promotes waste recycling, reduces resource extraction and lowering the carbon dioxide emissions [1–3]. The term “geopolymer” was sometimes used for AAM with a low calcium content [4], therefore relevant studies were also considered. The fly ash and GGBS based alkali-activated material (AAFS) could be cured in ambient conditions. The one-part AAM was also increasingly explored to achieve better workplace safety and convenience [5,6]. The AAFS had a similar mechanical behaviour to OPC concrete and could be predicted using existing OPC concrete models [7]. The AAFS has great thermal stability and chemical resistance [8]. The main reaction product from the ⁎

polymerisation reaction between the aluminosilicate binder and the sodium based activator is the sodium aluminium silicate hydrate (N-AS-H) gel, which remained chemically stable when exposed to sulfuric acid [9] and sodium chloride [10,11], making it an attractive option in aggressive environments. The use of fly ash could also reduce the drying shrinkage of alkali-activated slag (AAS) mixes [12]. Although the presence of GGBS made the mixture susceptible to decalcification compared to fly ash due to the formation of calcium silicate hydrate (CS-H) or calcium aluminate silicate hydrates (C-A-S-H), it significantly improved the pore refinement, limiting the ingress of deleterious ions [13]. The glass fibres and bars are innately corrosion resistant, which is crucial for extending the service life and reducing maintenance of the structures in harsh environments [14]. The GFRP bars are also lightweight (between 20 and 25% that of the density of steel), have high tensile strength (2–3 times the yield strength of steel) and are electromagnetically neutral [15]. However, instead of yielding, GFRP

Corresponding author. E-mail address: [email protected] (B. Yang).

https://doi.org/10.1016/j.engstruct.2019.109598 Received 15 April 2019; Received in revised form 27 August 2019; Accepted 27 August 2019 0141-0296/ © 2019 Elsevier Ltd. All rights reserved.

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and flexural loading. To improve the predictability and safety of circular GFRP-AAFS members, this study experimentally examined the structural behaviour of 14 GFRP-AAFS columns and beams with a practical height to diameter ratio (1150 mm in height and 215 mm in diameter; h/d = 5.3) and compared against 3 steel-AAFS concrete columns. This paper focused on the structural performances instead of the material development of AAM. The effect of reinforcement material, transverse reinforcement ratio (spiral pitch), longitudinal reinforcement ratio (number of longitudinal bars), and loading conditions was examined. Finally, the load-moment interaction diagrams were constructed using analytical methods and international design guidelines, and validated by the experimental results. A novel geometrical factor was introduced to enable the use of the methods in design guides specified for square or rectangular cross-sections to predict the behaviour of circular members. This simplified the design process and would be beneficial for the widespread use of GFRP-AAFS members in construction.

possesses a linear stress-strain relationship to the fracture failure in tension [16]. In compression, it is susceptible to the buckling failure due to its low elastic modulus [17]. Therefore, international design standards such as ACI 440.1R-06 [18] and CAN/CSA S806-12 [19] do not recommend the inclusion of GFRP bars in the compression capacity of the structure regardless its compressive strength. Several studies have been carried out on the behaviour of GFRP material under compressive loading. It was shown that the compressive strength was approximately half of the stated tensile strength and the elastic modulus in compressive was constant with that in tension [20]. The failure modes of the GFRP bars in compression were crushing (L/ db ≤ 7.3; unbraced length/diameter), buckling and crushing (7.3 < L/ db ≤ 14), and buckling (L/db > 14). Therefore, it is crucial to laterally support the longitudinal GFRP bars under compressive loads. In a complete GFRP reinforcement system, the lateral restraint to the longitudinal bars will be provided by GFRP transverse reinforcement, such as spirals [21,22] and stirrups [15,23]. It was reported that the GFRP hoops and spirals only reached 4–5% [24] of the ultimate tensile strain at the first peak. Despite of the low tensile strains, the GFRP stirrups had a tendency to open up prematurely and cause the collapse of the entire column [25]. This was attributed to the comparatively low stiffness at the bends. On the other hand, a uniform stress distribution could be achieved in the spirals or hoops, which allow the tensile strains to continue to develop until failure. Due to the low elastic modulus of GFRP, the reduction in the spacing between the transverse reinforcement would result in significant improvements in the confinement of the concrete core, compressive strains in the longitudinal GFRP bars, and avoiding early buckling [15,16,26,27]. With proper transverse reinforcement, the longitudinal GFRP bars contributed 5–11% to the load carrying capacity in axial compression [28–30]. As the cover concrete spalled, the contribution could increase to 23% of the second peak load [28]. While the majority of the studies focused on reinforced OPC concrete columns, the applicability of reinforced AAM columns has also been investigated. Sumajouw et al. [31] concluded that the current design provisions of the steel-reinforced OPC concrete could be adopted for AAM, by testing 12 slender fly ash based AAM columns reinforced with steel bars. The analytical method for OPC concrete columns could be used for steel-reinforced AAM columns with a suitable stress-strain relationship of AAFS [32]. In the case of GFRP reinforcement, the bond between GFRP bars and AAM was comparable to that of steel and AAM [33], and GFRP in OPC concrete [7]. The circular AAM columns fully reinforced with GFRP bars and hoops performed better than the OPC concrete counterparts [24]. The fully interaction diagrams of rectangular [34] and square [35] AAM columns reinforced with GFRP bars and stirrups were constructed. It was concluded that the AAM columns were more affected by the bending moment in the section, and a smaller rectangular stress block should be assumed as per [36]. Numerical analysis on rectangular and square GFRP-reinforced AAM members was carried out by incorporating confinement model and damage plasticity models [37,38]. Good agreements were found between the finite element analysis (FEA) results and experimental results. The literature review highlighted the potential use of GFRP-AAM composites in place of steel-OPC concrete composites in the construction industry. Currently, there is no study carried out for circular AAFS columns fully reinforced with GFRP bars and spirals under combined

2. Test program 2.1. Materials The GFRP bars and spirals were made from the pultrusion process, which pulled the uniformly aligned glass fibres through a resin bath, resulting in their high tensile strengths [39]. The 10 mm longitudinal bars had an ultimate tensile strength (ffu,bar) of 930 MPa, while the 8 mm GFRP spirals had a reduced ultimate tensile strength (ffu,spiral) of 650 MPa due to the bending process. The longitudinal ribbed steel bars were the 10 mm N10 (normal ductility deformed bar) bars, manufactured in accordance to AS/NZS 4671 [40]. The nominal tensile strength (ffu,steel) was 500 MPa, significantly lower than that of the GFRP. The properties of the GFRP and steel bars are reported in Table 1. The nominal areas reported by the manufacturers were used in the calculations. An equal-parts fly ash and GGBS mix was used in all specimens. The mix proportions are shown in Table 2. The oxide composition of fly ash and GGBS is shown in Table 3. The hydration moduli (HM = (Al2O3 + MgO + CaO)/SiO2; [41]) of fly ash and GGBS were 0.58 and 1.97, respectively, therefore GGBS was considered to be more active than fly ash in water. The GGBS with a greater hydraulic activity could promote the formation of C-S-H gels [9,42], which allowed the mix to be cured in ambient environment. The activator consisted of a 12 M sodium hydroxide solution and a commercially available sodium silicate solution (modulus Ms = SiO2/Na2O = 3.2, water content = 62% by mass), mixed at a 1:2.5 ratio. The molarity and mixing ratio were selected based on the findings and mixing procedures shown in [43–45]. The resultant Ms of the mixture was 1.4. The total binder consisted of fly ash, GGBS and the solid content in the activator, which amounted to 23% by mass of the mix. The Na2O in the activator was 5.1% the mass of the total binder, lower than the 5.7% limit recommended to avoid excess alkali thus efflorescence [46]. The total water content included the water in the activator, the water in the superplasticiser and the free water, which gave a total water to total binder ratio of 0.41. The aggregates complied with AS 2758.1 [47], which specified that natural aggregates had a water absorption of 2%. The measured 28-day compressive strength was 40 MPa.

Table 1 Properties of the reinforcement. Material

Bar type

Nominal diameter (mm)

Root diameter (mm)

Nominal area (mm2)

Tensile elastic modulus (GPa)

Yield strength (MPa)

Ultimate tensile strength (MPa)

Strain at ultimate tensile strength

GFRP

Longitudinal Spiral Longitudinal

10 8 10

9.2 7.4 –

66.5 40.7 78.5

59 55 200

– – 500

930 650 540

0.016 0.012 0.0027

Steel

2

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Table 2 Mix proportions. Constituents

Fly ash

GGBS

Activator

Fine aggregate

Coarse aggregate

Water

Superplasticiser

% by mass

10

10

8

21

45.7

4

0.3

2.2. Specimens

2.3. Test methods

The experimental program is shown in Table 4. A total of 4 parameters were varied, namely, the material of the longitudinal bar (steel or GFRP), the pitch of the GFRP spirals, the number of longitudinal GFRP bars, and the loading conditions (concentric, eccentric and flexural loading). The specimens were designated based on the 4 parameters. The first letter represented the material of the longitudinal bar, with “G” for “GFRP” or “S” for “steel”. The following number denoted the number of longitudinal bars in the cross-section. The spiral pitch in millimetre was then shown by the subsequent number and the final part denoted the loading condition, specifically, “C” for “concentric loading”, a number for load eccentricity in millimetre, or “F” for “flexural loading”. The 3–6 longitudinal bars corresponded to a reinforcement ratio of 0.55%, 0.73%, 0.92%, and 1.10%, respectively. The 40 mm, 80 mm and 120 mm spiral pitch was equivalent to a 2.75%, 1.39% and 0.94% transverse reinforcement ratio, respectively. The behaviour of the members loaded at three load eccentricities (e), low (25 mm), medium (50 mm) and high (75 mm), was studied. The reinforcement layout used in all specimens is shown in Fig. 1. A circular cross-section with a 215 mm diameter was adopted to facilitate the pouring of AAFS and ensure the specimen could be crushed using a 2000 kN universal testing machine. The specimens with an h/d ratio of 5.3 were considered as columns by design standards such as ACI 318 [48]. Meanwhile, the results were also representative as half-scale members. The unbraced effective length (Le) was the actual height of the specimens, given that the end connections were pinned connections. This resulted in a slenderness ratio (Le/r, r is the radius of gyration) of 21, which made the specimens short-unbraced columns (Le/r ≤ 22) according to AS 3600:2018 [49]. A concrete cover of 25 mm was deemed appropriate as the maximum aggregate size was 10 mm. To eliminate the end effects, a short 40 mm spiral pitch was used at the 200 mm segments from the top and bottom of the members. Two 30 mm strain gauges were attached to a pair of GFRP bars opposite to each other at the mid-height of the columns. An araldite seal was applied to prevent damage from casting and moisture ingress. The longitudinal GFRP bars extended from end to end in order to reveal the locations of the bars with the strain gauges. The desired spiral pitch was achieved by using 4 mm thick cable ties. The constructed reinforcement cage was placed in a PVC formwork before casting. The position of the cage was adjusted using two pairs of steel rods at top and bottom end regions of the member. The AAFS was synthesised by dry mixing at 60 revolutions per minute (rpm) for 3 min, followed by wet mixing at 60 rpm for 5 min. The AAFS was casted into the formworks alongside a number of 200 mm tall and 100 mm wide cylindrical moulds to determine the 28-day compressive strength. The curing of the cylinders and specimens took place in the ambient environment. All the specimens were demoulded after 24 h of curing.

The test setup is illustrated in Fig. 2. The specimens were held in place by a set of 150 mm tall steel collars fabricated using a 6 mm thick circular hollow section (CHS) with an internal diameter of 215 mm. For concentrically loaded columns, their two ends were in direct contact with the loading platens of the testing machine. This ensured no bending moment would develop from load eccentricity in the crosssection prior to the buckling failure. For eccentrically loaded columns, the pin supports were applied using two steel rollers welded to bottom plate of the collars. The eccentricity was measured from the centreline of the steel rollers to the centreline of the specimen. The universal testing machine was capable of applying a constant loading rate of 20 kN/min to the column specimens. The applied load and axial displacement from the machine were recorded, along with the strains in the strain gauges. Note that, the compressive strains are negative and tensile strains are positive unless specified. The reinforced beams were tested in a four-point bending setup. A pair of steel rollers were welded to a steel plate to transfer the load to the specimens. Another two steel rollers were placed at the two ends as supports. The distance between the four contact points was 317 mm, which left a 100 mm overhang at both ends. The same testing machine was set to exert a constant 2 kN/min to the specimen. In addition to the loads and strains, the data from a laser triangulator at mid-span that measured the transverse displacement of the tensile face was recorded. 3. Results and discussion 3.1. Test results of the concentrically loaded columns The axial load-axial deflection curves of the concentrically loaded columns are shown in Fig. 3, whereas the strains recorded in the longitudinal bars are shown in Fig. 4. The failure modes of some representative specimens are displayed in Fig. 5. The behaviour of G4-40-C and S4-40-C with a short spiral pitch of 40 mm is compared in this section. Immediately from Fig. 3, the peak load and stiffness in the elastic range of G4-40-C were slightly lower than those of S4-40-C, which was attributed to the greater stiffness of the steel bars in compression. At the peak load, the cover concrete started to spall when it reached its strain capacity, accompanied by a moderate drop in load soon after. The compressive load was then taken by the concrete core, which had a higher strain capacity than the cover concrete due to the confinement from the spirals. Both columns exhibited ductile behaviours after the peak load, which was indicated by the flat segment as shown in Fig. 3. The “steps” in the post-peak curves of both specimens corresponded to a fracture in the GFRP spirals. Once the column lost enough confinement from the spirals, it collapsed, as shown in large reduction in load at the end. The strains (Fig. 4) developed in the two bars on opposite sides were similar until the postpeak loading. This showed that the experimental setup was effective in

Table 3 Oxide analysis of the raw materials. Material

Al2O3

CaO

Fe2O3

K2O

MgO

MnO

Na2O

P2O5

SiO2

SO3

TiO2

LOI1

Hydration

Fly ash GGBS

% 23.9 13.1

% 7 43.2

% 7.9 0.8

% 1 0.3

% 1.3 5.5

% 0.1 0.2

% 0.4 0.3

% 0.5 0

% 55.9 31.4

% 0.3 4

% 1.3 0.6

% 0.4 0.6

modulus 0.58 1.97

1

LOI: loss on ignition 3

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Table 4 Experimental setup. Specimen

Longitudinal bar type

Number of longitudinal bars

Stirrup pitch (mm)

Loading condition

Longitudinal reinforcement ratio (%)

Transverse reinforcement ratio (%)

S4-120-C G3-120-C G4-120-C G5-120-C G6-120-C

Steel GFRP GFRP GFRP GFRP

4 3 4 5 6

120 120 120 120 120

C C C C C

0.86 0.55 0.73 0.92 1.10

0.94 0.94 0.94 0.94 0.94

S4-40-C G4-40-C G4-40–25 G4-40–50 G4-40–75 G4-40-F

Steel GFRP GFRP GFRP GFRP GFRP

4 4 4 4 4 4

40 40 40 40 40 40

C C 25 50 75 F

0.86 0.73 0.73 0.73 0.73 0.73

2.75 2.75 2.75 2.75 2.75 2.75

S4-80-C G4-80-C G4-80–25 G4-80–50 G4-80–75 G4-80-F

Steel GFRP GFRP GFRP GFRP GFRP

4 4 4 4 4 4

80 80 80 80 80 80

C C 25 50 75 F

0.86 0.73 0.73 0.73 0.73 0.73

1.39 1.39 1.39 1.39 1.39 1.39

Columns G4-80-C and S4-80-C behaved similarly until the peak load, with S4-80-C having a slightly higher peak load and greater stiffness, as shown in Fig. 3. In terms of the post-peak behaviour, the two columns were less ductile than their 40 mm spiral pitch counterparts. The adverse effect of doubling the spiral pitch on the post-peak behaviour was evident. Soon after the spalling of the AAFS cover, the GFRP spirals were subjected to nearly twice the tensile load compared to those with a halved spiral pitch. Fracture failure occurred earlier and was more destructive, leaving a much longer unconfined length of concrete and resulting in the collapse of the column. The strain in the steel bars reached εs, c = −0.004 at peak load as shown in Fig. 4. The lower strain indicated a lower degree of confinement to the concrete

eliminating the eccentricity in the axially loaded columns. The compressive strain in the GFRP bars reached εfrp, c = −0.006 at the peak load whereas the strain in the steel rebars reached εs, c = −0.007. The high strains indicated that effective confinement was achieved through the tightly spaced GFRP spirals. As shown in Fig. 5, the cover concrete of G4-40-C completed spalled at the mid-height, exposing the spirals and core concrete inside. The core concrete remained mostly intact owing to the closely spaced spirals. Fractures were found in the GFRP spirals, and the GFRP bars fractured at a similar location. The failure occurred on a horizontal plane, indicating a crushing failure in the concrete core. A similar failure mode was observed for S4-40-C, with the exception that the steel rebars buckled upon failure.

Fig. 1. Reinforcement details of the specimens (units in mm). 4

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Fig. 2. Test setup of columns and beams.

Four columns (G3-120-C, G4-120-C, G5-120-C and G6-120-C) with the same spiral pitch had a varying number of longitudinal bars. The large 120 mm spiral pitch was chosen to limit the effect of the transverse reinforcement. As seen in Fig. 3, in the elastic range, the different longitudinal reinforcement ratios did not correspond to a significant variance in stiffness. However, as the longitudinal reinforcement ratio increased, the specimens experienced an increasingly ductile post-peak response. G3-120-C and G4-120-C failed shortly after the peak load, whereas the “steps” were observed in the post-peak curve of G5-120-C and G6-120-C. This proved that the longitudinal reinforcement was able to improve the confinement effect of the transverse reinforcement. This was also evident from the strains recorded in the specimens (Fig. 4).

core. The data of G4-80-C was not recorded due to a malfunction in the datataker. Fig. 5 revealed that the shear failure occurred in the exposed region of the concrete core of both specimens after the spiral fractured. Rupturing of the GFRP bars and buckling of the steel bars contributed to the collapse of G4-80-C and S4-80-C, respectively. G4-120-C and S4-120-C showed a similar behaviour to G4-80-C and S4-80-C, however with significantly more brittle failure modes (Fig. 3). The load capacities of the two columns dropped to less than 10% peak load immediately after the peak load. Due to their low transverse reinforcement ratio, the strains in the longitudinal bars were also significantly lower, at εfrp, c = −0.003 and εs, c = −0.002, respectively (Fig. 4). 5

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The compressive strains recorded for the four columns at peak load were approximately εfrp, c = −0.002, −0.003, −0.004, and −0.005, respectively. Fig. 5 shows the failures in G3-120-C and G5-120-C. The concrete core was better confined in G5-120-C than G3-120-C, and the fracture in the longitudinal bar was less critical. In G5-120-C, fracture occurred in two bars in close proximity, which produced the two “steps” in load before the collapse. Therefore, a well confinement system involved both the longitudinal and transverse reinforcement. 3.2. Test results of the eccentrically loaded columns The axial load-axial deflection curves of the two groups of eccentrically loaded columns (one with 40 mm pitch and the other with 80 mm pitch) are shown in Fig. 6, whereas the strains recorded in the longitudinal bars are shown in Fig. 7. The failure mode of each specimen is displayed in Fig. 8. The first group with 40 mm spiral pitch is compared in this section. Compared to the concentrically loaded columns (Fig. 3), as the bending moment increased, the load capacities of the columns (Fig. 6) reduced significantly. Instead of vertical cracking around the perimeter of the column, horizontal cracks started to form at the end of the elastic range. Heavy spalling occurred when peak load was reached, leading to substantial reductions in axial load. The compressive stresses were sustained by the confined concrete and the tensile stresses were taken by the GFRP bars on the tension face, which was reflected by the flat portions in the post-peak curves. Notably, the residual load experienced by the three specimens (G4-40-25, G4-40-50 and G4-40-75) at this stage was comparable. This resulted in an increasingly ductile response as the eccentricity increased. This was consistent with the strains in the longitudinal bars as shown in Fig. 7. While the compressive strains at peak load decreased from εfrp, c = −0.003 in G4-40-25 to εfrp, c = −0.002 in G4-40-75, the tensile strains increased. As the load eccentricity increased, the compressive strains in the longitudinal bars decreased significantly to a level comparable to concrete’s strain at peak stress. Fig. 8 shows the failures in the specimens. Fractures in the bars on the compressive face were observed in all three specimens However, columns G4-40-75 lost a substantial amount of concrete in the core region due to the small area of AAFS that held the compressive stresses. The second group with 80 mm spiral pitch was notably less ductile than the first group. However, the same trend of the residual loads could be observed across the three specimens, as shown in Fig. 6. The differences between the two groups in peak load and residual load gradually diminished as the eccentricity increased. The strains in the bars (Fig. 7) were slightly lower than those achieved in the first group. Based on Fig. 8, a different failure mode was observed in G4-80-25 and G4-80-50. Note that the data on the compression bars in G4-80-50 was not available due to the malfunction of the datataker. The failure region was closer to the top than the other specimens, where the failure occurred at mid-height. This was attributed to the development of a plastic hinge, which described the plastic bending caused by the excessive deformation [50]. The instability in the reinforcement system caused localised plastic deformation. While the load capacities of the two columns were not noticeably affected, this may prove unsafe in real world applications. Therefore, a higher transverse reinforcement ratio was essential for uniformly distributing the stress along the entire unsupported length of the column. This phenomenon was documented by [34,35] for GFRP-reinforced square and rectangular columns.

(a) Concentrically loaded columns with 4 longitudinal GFRP bars

(b) Concentrically loaded columns with 120 mm spiral pitch

Fig. 3. Axial load-axial displacement curves for the concentrically loaded columns.

(a) Concentrically loaded columns with 4 longitudinal GFRP bars

3.3. Test results of beams (b) Concentrically loaded columns with 120 mm spiral pitch

The transverse load – transverse displacement curves of beams G440-F and G4-80-F are shown in Fig. 9. The load increased as a similar rate in the elastic range due to their longitudinal reinforcement layout. At this stage, tensile cracks developed on the tension face near the midspan and started to develop upwards. As the loading continued, shear

Fig. 4. Strains developed in the longitudinal bars of concentrically loaded columns.

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Fig. 5. Concentrically loaded columns after testing.

(a) 40 mm spiral pitch

(a) 40 mm spiral pitch

(b) 80 mm spiral pitch

(b) 80 mm spiral pitch

Fig. 7. Strains developed in the longitudinal bars of the eccentrically loaded columns.

Fig. 6. Axial load-axial displacement curves of the eccentrically loaded columns.

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Fig. 8. Eccentrically loaded columns after testing.

Fig. 11, a predominant vertical crack was observed in G4-40-F after testing. The GFRP bars ruptured due to the excessive bending in the specimen. Spalling occurred in the cover on the compression face. The failure mode of G4-80-F was similar due to their identical longitudinal reinforcement ratio. 3.4. Ductility The ductility of the members was quantified using a method proposed by Elchalakani et al. [15], which was given by Eq. (1).

DI =

ADE ABC

(1)

the ductility index (DI) measured the ratio of energy absorbed until failure (the point where load dropped below 85% peak load) to the energy absorbed to the end of the elastic range (the point before the peak where the load was 75% peak load). The energies were the area (ADE and ABC, respectively) under the curve up to the specific points. The method is demonstrated in Fig. 12. The DI of all the columns is reported in Table 5. Expectedly, a shorter spiral pitch increased the ductility of the specimens. The DI values of G4-40-C, G4-80-C and G4-120-C were 2.64, 2.07 and 1.80, respectively. By having a small spiral pitch, the failure was more gradual and each fracture in the stirrups was less devastating to the whole member. Secondly, the longitudinal GFRP bars outperformed the steel rebars in terms of DI. The DI of S4-40-C, S4-80-C and S4-120-C were 1.85, 1.74 and 1.59, respectively, which corresponded to an average reduction of 19% by using steel rebars. Notably, the most significant difference of 30% occurred with a small spiral pitch of 40 mm. As the transverse reinforcement decreased, the margin between the two materials decreased substantially. Therefore, it is essential to provide adequate transverse reinforcement to maximise the performance of the longitudinal GFRP bars. The longitudinal reinforcement also played an important role in the confinement, as shown in the comparison among G3-120-C, G4-120-C, G5-120-C and G6-120C. An overall increasing trend was observed in those four columns from 1.78 to 1.91. The longitudinal bars could improve the confinement provided by the transverse reinforcement, however the change in DI values was not as drastic as those by changing the spiral pitch. Finally,

Fig. 9. Load-displacement curves of GFRP-reinforced beams.

Fig. 10. Strains developed in the longitudinal bars of the beams.

cracks started to form. Both beams failed at a similar load. At peak load, the tensile strains reached a similar εfrp, t = 0.007 for both G4-40-F and G4-80-F (Fig. 10). The longitudinal bars on the compression face firstly developed compressive strains, then progressed to low tensile strains as the neutral axis moved closer to the compression face. The beams failed shortly after the peak load with a negligible residual strength. Based on 8

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Fig. 11. G4-40-F after testing. Table 5 Experimental results.

Fig. 12. The method used for calculating DI.

Specimen

Peak load (kN)

Bending moment (kN∙m)

Standard prediction of axial load capacity*

DI

S4-120-C G3-120-C G4-120-C G5-120-C G6-120-C

1385 1018 1179 1288 1381

– – – – –

1380 1226/1139 1223/1137 1221/1134 1218/1132

1.59 1.78 1.80 1.77 1.91

S4-40-C G4-40-C G4-40–25 G4-40–50 G4-40–75 G4-40-F

1529 1459 1037 523 318 127

– – 25.9 26.2 23.9 20.1

1380 1223/1137 – – – –

1.85 2.64 1.91 1.99 2.06 –

S4-80-C G4-80-C G4-80–25 G4-80–50 G4-80–75 G4-80-F

1433 1290 944 527 296 120

– – 23.6 26.4 22.2 19.0

1380 1223/1137 – – – –

1.74 2.07 1.59 1.96 2.21 –

the DI increased as the load eccentricity increased. The columns have an identical longitudinal reinforcement ratio on the tension face, which resulted in a comparable level of the residual loads.

* For columns with steel rebars as longitudinal reinforcement, the prediction was based on AS 3600:2018 [49]; For those with GFRP bars, the prediction was based on ACI 440.1R-15 [18]/ CAN/CSA S806-12 [19]

3.5. Load capacities

where Ag is the gross sectional area, As is the total area of the steel rebars, Afrp is the total area of GFRP bars, Es is the elastic modulus of steel and f’c is the 28-day concrete compressive strength. The coefficient α1 was always smaller than 0.85, making CAN/CSA S806-12 [19] the most conservative. Unlike AS 3600:2018 [49], the compressive strength of the GFRP bars were excluded as per ACI 440.1R-15 [18] and CAN/ CSA S806-12 [19]. The results are reported in Table 5. Note that the results were identical for columns with various spiral pitches as confinement was not considered in those models. For the columns with steel rebars, S4-40-C, S4-80-C and S4-120-C were 11%, 4% and 0% higher than the predictions by AS 3600:2018 [49]. This indicated that the design of transverse GFRP reinforcement in steel reinforced concrete members could be carried out safely given that a sufficient transverse reinforcement was provided. For GFRP bars, G4-40-C, G4-80-C and G4-120-C were 19%, 5% and −4% higher than the standard predictions by ACI 440.1R-15 [18] and 28%, 13% and 4% higher than CAN/CSA S806-12 [19]. The larger increment over the

The load capacities of the concentrically loaded columns were compared against the predictions from various international design guidelines for steel and GFRP reinforcements (Eqs. (2)–(4)). AS 3600:2018 [49] was the Australian standard for steel reinforcement, whereas ACI 440.1R-15 [18] and CAN/CSA S806-12 [19] were the U.S. and Canadian standards for GFRP reinforcement. The proposed guideline for geopolymer, referred to as “Geopolymer Handbook” [36] in this study, adopted a similar method to AS 3600:2018 [49] for axially loaded steel-reinforced AAFS columns.

PAS; GeopolymerHandbook = 0.85f c' (Ag − As ) + 0.0025Es As

(2)

PACI = 0.85f c' (Ag − Afrp )

(3)

PCAN / CSA = α1 f c' (Ag − Afrp ); α1 = 0.85 − 0.0015f c' (≥0.67)

(4) 9

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In the second method, the rectangular stress block (RSB) methods from four international design guidelines [18,19,36,49] were adopted and modified for circular sections. The RSB was already an approximation of the integration of the concrete in compression, however it only worked under the assumption of a uniform width throughout the section. For a circular section, this would lead to discrepancies. However, this method remained easy-to-use and widely adopted for designs of rectangular or square sections, therefore a factor that improved the accuracy of approximation would benefit the designs of circular sections. In those standards, a factor γ or its equivalent controlled the height of the RSB, whereas α2 controlled the concrete compressive strength. The values of the two factors for the AAFS (f′c = 40 MPa) from various standards are reported in Table 6. A factor βc was introduced to modify α2 in order to achieve a better representation of the bending moment in the RSB. This factor is the ratio of the internal moment of the concrete in compression Mc,circular to the moment obtained using the unmodified RSB method, Mc,RSB, both taken at the extreme compressive fibre. The ultimate moment of the section was taken with respect to the neutral axis. The full derivation is illustrated in Fig. 14, and shown as Eqs. (5)–(12). Note that for OPC specimens, Eq. (5) was adopted from [51].

standard predictions was due to the more conservative nature of two standards, which did not include the compression capacity of the GFRP bars. With a very low transverse reinforcement ratio (spiral pitch of 120 mm), the unfactored predictions by ACI 440.1R-15 [18] was unsafe. Therefore, the columns must be adequately transversely reinforced to ensure safe designs. By comparing the two materials, the steel rebars outperformed the GFRP bars by 11% on average. The difference increased from 5% with a 40 mm spiral pitch to 17% with a 120 mm pitch, which showed that the GFRP bars could be used as an alternative to steel rebars given that adequate transverse reinforcement was provided. For the columns with varying number of longitudinal GFRP bars, the predictions by ACI 440.1R-15 [18] started from over-predicting by 17% for G3-120-C to under-predicting by 13% for G6-120-C. This again showed that with inadequate confinement, the GFRP-AAFS system was not safe in terms of the standard predictions. This could be mitigated by having a larger longitudinal reinforcement ratio and/or a larger transverse reinforcement ratio. The differences in the load capacities showed that the GFRP bars contributed to the compression capacity of the member, therefore it would be overly conservative to use the standards for those with larger reinforcement ratios. Lastly, the axial load capacities of the eccentrically loaded columns with a 40 mm spiral pitch were greater than those with an 80 mm spiral pitch, similar to their concentrically loaded counterparts. The area of concrete in compression decreased, therefore the effect of confinement decreased. As the load eccentricity increased, the difference reduced from 10% at e = 25 mm to 7% at e = 75 mm. Additionally, the difference in the transverse load capacities of AAFS beams G4-40-F and G480-F was only 6%.

σc =

Ec =

⎧ 0whenεc ≤ 0; Ec εc whenεc > 0. ⎨ 1 + εc' ⎩ εc 2f c' εc'

(6)

(εtop − εbot ) y

εc = εtop −

2r

3.6. Interaction diagrams

dn =

Two methods were used to construct the interaction diagrams of the columns. The first method involved the use of a theoretical approximation method that divided the section into a finite number of strips for analysis. The second method was derived from the methods used for constructing interaction diagrams for square or rectangular sections in the international design guidelines [18,19,36,49]. A factor was introduced to make them compatible with circular sections. An analytical approach was used to obtain the first peak load and its corresponding moment at any given eccentricity in the first method. The section was divided into 40 strips, and the strain at each strip was determined based on the strain at the extreme compressive fibre of AAFS and the GFRP bars at the bottom. A strain of εc = −0.003 was assumed at the top and the stain in the bottom bars varied from εfrp,c = −0.003 to εfrp,t = 0.016, which corresponded to the ultimate tensile strain of the GFRP bars [39]. The Desayi and Krishnan [51] concrete strain-strain model was used to determine the average unconfined stress in each strip. The model was originally proposed for OPC concrete, however was proved to be appropriate for a similar equal-parts fly ash and GGBS based mix [7]. The strain at peak stress in AAFS was assumed to be εc = −0.002, recommended by Mander et al. [52] in analytical analysis and proved by the experimental data by Dong et al. [7]. The Kappos and Konstantinidis [53] confinement model was used to obtain the confined stress in the effectively confined region of each strip. Based on Mander et al. [52], at half-way between two hoops, the concrete received the least amount of confinement, making it the most critical. The border of the effectively confinement region in the longitudinal plane followed an arc bound by 45˚ tangents to the horizontal planes. This was considered a valid approximation for spiral reinforcement. A linear stress-strain relationship with a constant elastic modulus in compression and tension was adopted for the GFRP. The force and the bending moment with respect to the extreme compressive fibre of the section were determined in each strip. The summation of the forces and moments was the section load and moment capacities, respectively. This method is demonstrated in Fig. 13.

(5)

where0 ≤ y ≤ 2r

(7)

2rεtop εtop − εbot

lchord = 2 r 2 − (r − y )2

(8) (9)

dn

Mc, circular =

∫ y∙lchord ∙σc dy 0

(10)

γ ∙ dn

Mc, RSB =

∫ y∙lchord ∙α2 ∙fc' dy 0

βc =

Mc, circular Mc, RSB

(11)

(12)

where Ec is the elastic modulus of concrete, given by Eq. (6). The strain at peak stress εc' was assumed to be −0.002 [52]. The strain εc linearly varied in the section with respect to a given depth of y from the top, based on the strains in the extreme compressive fibre of concrete (εtop = −0.003) and strains in the bottom longitudinal bars (εbot = −0.003 to 0.016). The depth of the neutral axis is dn. The radius of the section is r, which equalled 107.5 mm, and the chord length lchord is a function of y. Therefore, based on Eqs. (10) and (11), βc is a function of γ, which varied in different design guidelines. The integrals were solved numerically, and it was found that βc gradually reduced as the neutral axis shifted upwards. However, it tended to converge at a minimum value, which could be used conservatively for design. A larger γ corresponded to a reduced βc. The βc value is reported in Table 6 for each standard. The γ in Geopolymer Handbook [36] was 0.65, notably smaller than the other OPC concrete standards. This meant the RSB in AAFS was smaller than OPC concrete. The βc of 0.95 was much greater than those of the other standards, especially CAN/ CSA S806-12 [19], with a γ of 0.87 and βc of 0.85. The results obtained in this study were compared against a similar experimental study using OPC concrete [54]. The members had a 215 mm diameter cross-section and 4 longitudinal GFRP bars restrained 10

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Fig. 13. The theoretical prediction of the force and moment at the cross-section of the GFRP-reinforced members.

effectiveness of the model, the moment was normalised by d3f c' and the load was normalised by d 2f c' , where d is the diameter of the section. In this figure, it is evident that the confinement provided by the GFRP bars and stirrups between the two concretes was comparable for the concentrically loaded columns. As the eccentricity increased, the difference between them considerably increased, especially for the columns loaded at e = 50 and 75 mm and the beams. This showed that the AAFS members had a smaller stress block than their OPC concrete counterparts, which resulted in reduced moment capacities at higher eccentricities. The predictions based on AAFS and OPC concrete were similar once normalised, and the difference between the predicted curves of 40 and 80 spiral pitch was smaller than those obtained in the experiment. This was caused by the conservative estimate of the effectively

Table 6 The comparison of various factors used in the RSB method (f’c = 40 MPa). Standard

α2

γ

βc

AS 3600:2018 [49] ACI 440.1R-15 [18] CAN/CSA S806-12 [19] Geopolymer handbook [36]

0.79 0.85 0.79 0.85

0.87 0.77 0.87 0.65

0.85 0.93 0.85 0.95

by GFRP spirals at a 40 or 80 mm pitch. The interaction diagrams constructed using the first method are shown in Fig. 15. The OPC concrete had a compressive strength of 34 MPa, lower than the AAFS (40 MPa) used in this study. Therefore, in order to compare the

Fig. 14. Obtaining βc. 11

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AAFS using the CAN/CSA S806-12 [19] was unsafe, however the proposed curve showed good agreements with the experimental results. This design standard was relatively conservative, therefore the discrepancy for the concentrically loaded AAFS columns was much larger. When used for OPC concrete members, it considerably underpredicted the load and moment capacities. Another design standard for GFRP, the ACI440.1R-15 [18] shown in Fig. 16(b), was less conservative, therefore its corresponding modified calculation was more accurate and relatively safe. It could be seen that a capacity factor should be used in the proposed curve as the safety margin was almost zero for the specimens loaded at high eccentricities. The AS 3600:2018 [49] [Fig. 16(c)] for steel reinforcement was suitable for predicting the capacities of GFRP members after adopting the proposed modifications. However, it was conservative for OPC concrete members. The Geopolymer Handbook [36] was the only design guideline specifically made for AAFS [Fig. 16(d)], which took the reduced stress block into account. It could be seen that it was the most appropriate for AAFS members, while being overly conservative for OPC concrete members. Its proposed modification showed the best agreements to the experimental results among the four standards. Overall, it is still recommended to use a capacity factor with caution when designing for GFRP-reinforced AAFS members. The proposed factor βc proved to be safe and convenient for designing circular sections.

Fig. 15. Theoretical predictions of members constructed with AAFS (this study) and OPC concrete [54]. Note that “40” and “80” refer to the spiral pitch.

confinement area. It was shown in this study that the longitudinal bars also contribute to the confinement of the concrete, therefore the reduction in the effectively confinement area in the longitudinal direction proposed by Mander et al. [52] could be relaxed, given that a sufficient number of longitudinal bars was in place. The predictions of the OPC concrete members were conservative, whereas the predictions became unsafe for some AAFS members with large bending moments in the cross-section. Therefore, it is advised that a safety factor should be used for the theoretical predictions of the AAFS members. The interaction diagrams constructed using the second method are shown in Fig. 16. In this figure, the unedited standard predictions [18,19,36,49] are presented alongside the model proposed in this study, which included the contribution of the GFRP bars in compression and the factor βc for the circular cross-sections. It could be seen that overall, the proposed model was more accurate and safer than the standard predictions. In Fig. 16(a), the predicted interaction diagram of

4. Conclusions The following conclusions could be drawn in this experimental study of 17 reinforced AAFS members (14 with longitudinal GFRP bars and 3 with steel bars): 1. A large transverse reinforcement ratio could improve the load and moment capacity, ductility and the failure mode. The columns with a 40 mm spiral pitch were 7% and 11% greater than their 80 mm spiral pitch counterparts in terms of load capacity and ductility,

(a) CAN/CSA S806-12

(c) AS3600:2018

(d) Geopolymer Handbook

(b) ACI440.1R-15

Fig. 16. Original and modified standard predictions of members constructed with AAFS (this study) and OPC concrete [54] 12

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2.

3.

4.

5.

respectively. With a sufficiently high transverse reinforcement ratio, it was recommended to include GFRP in axial load capacity calculations. The columns with a larger number of longitudinal GFRP bars had improved load capacity and more importantly, ductility. The longitudinal bars could improve the confinement effect of the transverse reinforcement, which may be incorporated in the confinement models to better define the effectively confined regions. The columns with the GFRP longitudinal bars had slightly reduced load capacities (10%) but increased ductility (24%) on average than those with longitudinal steel rebars. The confinement effect of the transverse reinforcement decreased as the load eccentricity increased due to the reduction in the area of concrete in compression. The residual strengths of columns loaded at different eccentricities were comparable, resulting in increased ductility indices as the eccentricity increased. A new geometrical factor βc was introduced in the current design rules in various design guidelines to construct interaction diagrams for circular columns. The results had shown that conservative yet accurate predictions could be achieved using the factor. It was evident that AAFS members had a smaller stress block than those specified in OPC concrete standards. The proposed method would facilitate the design of the circular GFRP-reinforced AAFS members, which could be effectively used in corrosive environments, for example, as bridge piers in coastal zones, in cement plants and as underground sewer pipes.

[11] [12]

[13]

[14]

[15]

[16]

[17]

[18]

[19] [20]

[21]

Declaration of Competing Interest

[22]

We confirm that there are no known conflicts of interest associated with this publication and there has been no significant financial support for this work that could have influenced its outcome.

[23]

[24]

Acknowledgements This research is supported by 111 Project of China (Grant No. B18062). The authors are grateful for the donations and support provided by Pultron Composites, New Zealand and Anthony Miles from Sika Australia. Appreciation is also given to former student Mr Manuj Sharma and Mr Mariwais Nazifa, laboratory technicians Mr Jim Waters, Mr Brad Rose and Mr Matt Arpin for their help with the many practical aspects of this project.

[25]

[26]

[27]

[28]

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