A solution for sea-sand reinforced concrete beams

Mei-ni SU1, Liang-liang WEI2, Zhi-wen ZENG3, Tamon Ueda4, Feng XING5, Ji-Hua ZHU6*,

1 Lecturer, School of Mechanical, Aerospace and Civil Engineering, University of Manchester, Manchester, M1 3NJ, UK2 PhD Candidate, Laboratory of Engineering for Maintenance System, College of Engineering, Hokkaido Univ., Sapporo 060-8628, Japan.3 M.Sc Candidate, Guangdong Province Key Laboratory of Durability for Marine Civil Engineering, School of Civil Engineering, Shenzhen University, Shenzhen, Guangdong 518060, PR China.4 Professor, Laboratory of Engineering for Maintenance System, Faculty of Engineering, Hokkaido Univ., Sapporo 060-8628, Japan.5 Professor, Guangdong Province Key Laboratory of Durability for Marine Civil Engineering, School of Civil Engineering, Shenzhen University, Shenzhen, Guangdong 518060, PR China. 6 Professor, Guangdong Province Key Laboratory of Durability for Marine Civil Engineering, School of Civil Engineering, Shenzhen University, Shenzhen, Guangdong 518060, PR China. (Corresponding author: zhujh@szu.edu.cn)

Abstract：The popularity of reinforced concrete (RC) structures leads to increasing

demand for sands, cement, aggregates and other raw materials. In the recent decades,

river sand has been used to replace sea-in order to solve the resource shortages in

many countries. However, sea-sand concrete might cause corrosion of steel re-bars

and result in structure deterioration. Impressed current cathodic protection (ICCP) is

an efficient method to prevent corrosion of re-bars, while bonding carbon fibre mesh

to the RC structures can help improve the loading capacity of the deteriorated

structures. This study proposes a new dual-functional intervention method, the

impressed current cathodic protection – structural strengthening (ICCP-SS) method,

to retrofit the deteriorated sea-sand RC structures by using the carbon - fabric

reinforced cementitious matrix (C-FRCM). The C-FRCM composite, comprised of

carbon fabric mesh and inorganic cementitious matrix, is both the anodic material for

Su, M.N., Wei, L.L., Zeng Z.W., Ueda, T., Xing, F., Zhu, J.H. “A solution for sea-sand reinforced concrete beams”, Construction and Building Materials 204, 586-596.

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ICCP and the structural strengthening material. This paper presents an experimental

program consisting of 11 simply supported beams, 9 of which were casted by

simulated sea-sand and subjected to accelerated corrosion for 130 days. The

specimens casted with simulated sea-sand were afterwards externally bonded with C-

FRCM composite, exposed to ICCP for another 130 days, and finally tested. In this

study, the loading capacity and deflection at midspan of the beams, as well as the

open circuit potential (OCP) of re-bars were measured to assess the effectiveness of

the intervention method. The proposed method has been shown to be effective in

retarding the corrosion of steel re-bars and improving the loading capacity of the

corroded specimens. In addition, this paper compares the experimental results with

the capacity predictions set out in ACI 440.2R-08 for FRP strengthening system and

ACI 549.4R-13 for FRCM strengthening system, which have been found to be rather

conservative for the flexural design of retrofitted beams.

Keywords： C-FRCM; corrosion; impressed current cathodic protection (ICCP);

reinforced concrete; sea sand; simply supported beams; structural strengthening.

1 Introduction

RC structures are widely popular in the construction industry. However, the huge

demand of concrete is resulting in the resource shortages, such as fresh water and

river sand. Nowadays, sea sand has to be used when river sand is unavailable [1,2].

However, sea sand normally contains high percentage of chlorides, which could be up

to 2% (i.e. the known Cl- concentration in sea water) [3]. In order to avoid the

corrosion of steel re-bars, it should be thoroughly washed before being used. The ACI

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201.2R-2008 [4] gives limitations of 0.06% and 0.08% (by the weight of cement,

similarly hereinafter) for reinforced concrete in moist environments with and without

exposure to external chlorides, respectively. BS EN 206-1:2000 [5] specifies the

limits of 0.1% and 0.4% in concrete containing prestressing steel reinforcement and

steel reinforcement or other embedded metal, respectively. In China, the figures are

limited to 0.06% and 0.10% for sea sand concrete in moist environments with and

without exposure to external chlorides, respectively; it goes up to 0.3% in dry

conditions and plain concrete [6]. However, washing sea sand will cost a great amount

of fresh water, electricity and human resources. What is worse, it is not easy to

guarantee the chloride contents in the sea-sand, and if the target chloride content is not

satisfied, sea sand concrete might cause the corrosion of steel re-bars and the

degradation of RC structures. The deterioration of RC structures might result in

enormous economic loss and, more importantly, the loss of safety.

Impressed current cathodic protection (ICCP) is a technique for retarding the

further corrosion of metals [7]. It has been used to protect the steel re-bars in

structures that use RC since the 1970s [8]. In the process of ICCP, an impressed

current is applied to the steel reinforcement to charge the steel negatively. As a result

of the cathodic polarization, the steel becomes cathode and corrosion is impossible

[9]. Many studies have investigated the effectiveness of ICCP systems on steel

protection. These studies have focused on interrupted ICCP [10], criteria and

important parameters of ICCP [11], RC structures in a marine environment [12], and

persistent protective effects of field structures [9]. The sound effect of ICCP technique

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has been proven in the literature [9-11].

A widely-used method of strengthening a degraded RC structure is to externally

bond it with strengthening materials such as steel plates, or FRP (fibre reinforced

polymer) plates/sheet/meshes [13,14]. Comprehensive experimental [15–18] and

numerical investigations [19–21] have been conducted in past decades to investigate

the optimum design method, the practical construction procedure, and the key factors

of the strengthening method using externally bonded FRP. Analytical models were

also studied by a great number of researchers regarding design models and failure

modes [22,23]. There are also a few design codes for FRP-strengthened RC

structures, such as the JSCE [24], the fib bulletin 14 [25], the ACI guide [26], and the

ISIS Canada Design Manual [27]. To date, the FRP strengthening technique and

corresponding design methods have been well developed.

On the one hand, using the ICCP system can efficiently impede the ongoing

corrosion of re-bars in sea-sand RC structures, but it cannot recover the strength loss

due to the corrosion at an early stage. On the other hand, externally bonded FRP is a

widely used retrofitting method to improve the loading capacity of degraded RC

structures, but it cannot impede the further corrosion of re-bars. Therefore, this study

proposed a novel retrofitting method by taking advantage of the both techniques,

termed as ICCP-SS (impressed current cathodic protection – structural strengthening)

method. In this new intervention method, a dual-functional carbon-fabric reinforced

cementitious matrix (C-FRCM) was used as both the anode material in the ICCP

system and the strengthening material in the SS system [28]. The C-FRCM composite

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was comprised of carbon fabric mesh and inorganic cement-based matrix. The carbon

fabric mesh embedded in the C-FRCM has a superior conductivity characteristic with

a reasonable and acceptable cost in comparison with normal anode material such as

costly titanium mesh, platinum, niobium and mixed metal oxide (MMO), meanwhile,

it also has an advanced light weight – high strength and a strong resistance to

corrosion characteristics comparing with traditional steel plate for strengthening.

Though there are a great number of studies on both the ICCP and SS techniques, the

ICCP-SS method is relatively new with limited investigations [29-31]. The effects of

applied current on the C-FRCM and the bonding behaviour, which might cause a

negative effect on the SS system, still need further careful investigation.

This paper presents a series of simulated sea-sand simply supported beam tests.

To measure the diverse corrosive effects, a total of 11 concrete specimens were cast, 9

of them with an amount of NaCl to simulate the sea-sand concrete. After curing, the

specimen experienced 130-day accelerated corrosion and 130-day cathodic protection.

Test results were recorded and compared to assess the effectiveness of the ICCP

technique, the SS technique and the ICCP-SS technique on the corroded beams. The

American Concrete Institute’s design guidelines, ACI 440.2R-08 [26], Guide for the

Design and Construction of Externally Bonded FRP Systems for Strengthening

Concrete Structures, and ACI 549.4R-13 [32], Guide to Design and Construction of

Externally Bonded Fabric-Reinforced Cementitious Matrix (FRCM) System for

Repair and Strengthening Concrete and Masonry Structures, were used to predict the

design capacities of the tested beams, which were then compared with the test results.

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The applicability of the existing design code - ACI 440.2R-08 [26] and ACI 549.4R-

13 [32] for the sea-sand RC structures repaired by ICCP-SS technique is also

discussed.

2 Experimental programme

An experimental program including 11 simulated sea-sand reinforced concrete (RC)

beams was carried out in the structural laboratory of Shenzhen University.

2.1 Test specimens

The specimens were designed to be repaired by ICCP, SS or ICCP-SS techniques. All

the control specimens experienced an accelerated corrosion process before bonding

the C-FRCM composite onto the soffit. Afterwards, different constant currents were

applied to the specimens which are designed to be repaired by ICCP and ICCP-SS

techniques, whereas it is not needed for specimens repaired by SS technique. For

ICCP specimens, the C-FRCM composite was removed after the ICCP before four-

point bending tests.

Following the abovementioned process, eleven test specimens were divided into

five groups: (1) two specimens without NaCl (i.e. reference specimens); (2) one

specimen contained NaCl without any repairing techniques (i.e. reference specimen);

(3) one specimen contained NaCl and was repaired by SS technique; (4) four

specimens contained NaCl and were repaired by ICCP technique; (5) three specimens

contained NaCl and were repaired by ICCP-SS technique. The weight of NaCl was

3% of the cement and was contained in the mix of concrete. After the curing period,

the specimens were exposed to accelerated corrosion, followed by the ICCP process.

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The labelling system used for the specimens is given in Table 1. The detailed design

of the beam specimens is shown in Fig. 1. If a test is repeated, a letter “R” is added in

the label of the specimen.

2.2 Properties of the material

Table 2 shows the concrete mix proportions. The average compressive strength of the

concrete cubes in 150 mm was found to be 53 MPa. The nominal diameter of re-bars

used in the specimens was 10-mm. Table 3 shows the mix proportions of the

inorganic cementitious matrix. The material properties of steel re-bars, carbon fabric

mesh and cementitious matrix were obtained through tests according to the ASTM

E8/E8M [33], the ASTM D4018 [34], the BS EN 196-1 [35], as presented in Table 4.

The material properties of the C-FRCM composite was determined by the

uniaxial tensile tests according to AC434 [36] (see Fig. 2). The failure mode was

slippage of the carbon fabric within matrix after initially cracking of matrix, which is

a combination of pull-out failure and tensile fracture failure. A typical stress - strain

curve of C-FRCM from tensile test is shown in Fig. 3, together with a typical stress-

strain curve of one bundle of carbon fibre impregnating with epoxy resin. The tensile

behaviour could be characterized as a bilinear curve, in which the first phase

represents the uncracked behaviour of C-FRCM, while the second part of the curve

indicates the behaviour after the occurrence of cracks on the cementitious matrix.

According to the recommendation in AC434 [36], the tensile modulus of elasticity of

C-FRCM is defined by two points on the second part of the curve, at a stress level

equal to 0.60ffu and 0.90ffu, as given in Eq. (1), which was calculated as 195 GPa. The

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peak tensile stress and the corresponding strain were found to be 1584 MPa and

1.04%, respectively. The tensile modulus of elasticity of carbon fibre was found to be

223 GPa, and the ultimate tensile stress and strain were 3519 MPa and 1.58%,

respectively. The material properties of C-FRCM and carbon fibres are shown in

Table 4.

(1)where,

Efrcm = tensile modulus of elasticity of cracked C-FRCM;

ffu = ultimate tensile strength of C-FRCM;

= tensile strain of C-FRCM at a stress level equal to 0.6ffu;

= tensile strain of C-FRCM at a stress level equal to 0.9ffu.

2.3 Accelerated corrosion procedure

An amount of NaCl about 3% by the weight of cement was added to the concrete

mixture to simulate the sea-sand concrete. This amount of chloride, which was greater

than any chloride threshold value for corrosion onset reported in the literature [3], de-

passivated the re-bars and induced corrosion. No NaCl was included in the concrete

mixture that was used to produce the reference specimens. In order to induce

corrosion damage in the tested specimens within a reasonable period, the specimens

were placed in an open-air space and were subjected to two wet–dry cycles per week

(2.5 days wet followed by 1 day dry) continuing 130 days.

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2.4 Installation of C-FRCM to the soffit of RC beams

The C-FRCM composite was installed on the soffit of corroded RC beams after

accelerated corrosion period. Firstly, the weak segment at the soffit surface of beams

was polished away to exposure the harder coarse aggregate and increase the surface

roughness. This is done to improve the bonding performance between C-FRCM and

substrate concrete. The treated soffit of beams should be kept saturated for 12 hours

before bonding the C-FRCM. The first layer of 5 mm cementitious matrix was applied

to the treated soffit surface of the beam. Afterwards, the carbon fabric mesh was laid

on the top of cementitious matrix and was impregnated into the matrix gently. Finally,

the second layer of 5 mm cementitious matrix was applied to cover the carbon fabric

mesh. The nominal thickness of C-FRCM was therefore approximately 10 mm. The

bonding area is the full size of the soffit side of the beam.

2.5 ICCP process

The ICCP was applied in 28 days after installing the C-FRCM composites. The re-

bars were connected to the negative terminal and the carbon fabric mesh embedded

into the C-FRCM acted as anode was connected to the positive terminal of an external

multi-channel DC power supply to apply protective current to the corroded steel re-

bars (Fig. 4). The ICCP system was operated in the open air space for 130 days. Two

applied current densities were employed - 26 mA/m2 (small current density) and 80

mA/m2 (large current density) of the re-bars’ surface area. The currents were

measured and the open circuit potential (OCP) values of the embedded steel were

recorded at the interval of ten days.

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The specimens were monitored for corrosion activity with internal reference

electrode (RE) and external instrumentation. Before the concrete was cast, a RE was

placed vertically on the upper side of the mid-span of each beam during assembly of

the steel cage. The embedded REs were calomel RE saturated by KCl solution. The

saturated KCl solution was kept inside the probe by a rubber cap. The measurements

were conducted in light of the requirements of ASTM C876-09 [37].

2.6 Four-point bending tests

A four-point loading set-up (Fig. 5) with a hydraulic jack was used to test the

specimens. The applied loads were measured by a load cell placed between the

hydraulic jack and distribution of beam. Deflection at the constant moment region of

specimens was measured by three LVDTs. The specimens were loaded by

displacement control at a loading rate of 0.5 mm/min. A computer-based data

acquisition system recorded the data at a frequency of 1Hz.

3 Experimental results and discussions

3.1 Results and discussions on the ICCP performance

During the 130-day operation of the ICCP, the OCP values of the re-bars

of seven selected specimens were recorded and plotted in Fig. 6. In

accordance with the recommendations of ASTM C876-09 [37], if the

OCP value is greater than -126 mV, it indicates that the embedded

steel has only 10% possibility of being corroded; if the OCP value is

less than -275 mV, it demonstrates that the embedded steel has

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90% possibility being corroded; if the OCP value is between these two

values, it means the status of the re-bars is uncertain. The

potentials of all specimens showed an increasing trend. It might be

because the potentials were recorded in summer time, which is

relatively dry during that period. It is reasonable that the potentials

were fluctuated to some extent due to the changes of local climate.

The potential of the re-bars in the reference beam without NaCl

(specimen SB) is above the -126 mV level during the whole

monitoring period. The specimens with NaCl were generally below

the line of -275 mV. As for the specimens that contained NaCl but

hadn’t been protected by ICCP, they stayed below the line of -275

mV, since the steel re-bars embedded in SB-C and SB-C-F1

specimens were subjected to corrosion continuously. Please note

that the decreasing trend of the potentials of these two specimens

after 160 days could be due to the season changes. However, when

the ICCP starts to operate, the potential increases and gets closer to

the margin of -126 mV as the time goes by. If greater current

densities had been applied to the specimens, the protection effects

may be more obvious. However, it should be noted that too great a

current density would also result in the premature deterioration of

the bond interface. In some investigations [29], higher applied

current densities were adopted from 125 to 200 mA/m2 of steel

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surface area. The electrochemical parameters measured in this

study also indicated that the embedded steel has a high chance of

being successfully protected. However, an amount of gaseous

yellow liquid appeared on the surface of the carbon fabric after 474

hours in the process of ICCP, which may cause the separation of the

carbon fabric from the concrete interface [29]. Therefore, smaller

values of applied current densities (26 and 80 mA/m2) were chosen

in this study.

3.2 Results and discussions on the loading responses

The test results of ultimate loads are presented in Table 5, while

the typical failure modes of the tested beams are shown in Fig. 7.

During testing, the first major flexural crack occurs in the constant

moment region, followed by some minor shear cracks as the load

increases. All the tested beams had the re-bars yielding, and finally

failed upon the carbon fibre meshes slippage and fracture at the

matrix cracked section and the compression concrete crushed. No

delamination between C-FRCM composite and concrete was

observed.

Fig. 8(a) shows the comparison of flexural response of

specimens SB, SB-R, SB-C and SB-C-F1, indicating the structural

strengthening effect provided by C-FRCM composite. For the

reference beams (SB and SB-R), the loading capacities are 50.4 kN

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and 47.2 kN (average load = 48.8 kN). For the corroded reference

specimen without ICCP (SB-C), the load capacity is 33.4 kN, which is

31.5% lower than the average load capacity of the reference beams

(SB and SB-R). This might be attributed to the reduction in the

effective area of re-bars. The ultimate capacity of the beam

strengthened with C-FRCM without ICCP (SB-C-F1) is 43.2 kN, which

is 29.3% higher than the average ultimate capacity of the un-

strengthened beam (SB-C). This demonstrates that bonded C-FRCM

can effectively improve the flexural capacity of corroded beams.

However, it is still lower than the flexural capacity of the reference

beams (SB and SB-R), possibly due to the insufficient strengthening

material.

A total of 4 beams were protected by sole ICCP after accelerated

corrosion. Fig. 8(b) shows the flexural behaviours of reference and

control beams, which can assess the effectiveness of cathodic

protection using C-FRCM composite as anode. The flexural

capacities of these beams were found to be 42.3 kN, 46.9 kN, 47.0

kN and 43.2 kN for SB-C-IS, SB-C-IS-R, SB-C-IL, and SB-C-IL-R,

respectively, which are 26.6%-40.7% higher than the un-repaired

beam (SB-C), but 3.7% - 13.3% lower than the reference beams (SB

and SB-R). On the one hand, it demonstrates that the operation of

ICCP technique can effectively impede the further corrosion of re-

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bars, so that the specimens repaired by ICCP have higher loading

capacities than the specimens without the operation of ICCP due to

the differences in effective cross-section of re-bars and bonding

performance. On the other hand, it shows that ICCP treatment

cannot help to recover the loading capacities due to the existed

corrosion of re-bars.

Fig. 8(c) shows the loading response of reference and control

beams to show the benefits of ICCP-SS intervention method. The

ultimate capacities of the beams retrofitted by ICCP-SS method (SB-

C-F1-IS, SB-C-F1-IS-R and SB-C-F1-IL) were found to be 41.4 kN, 44.0

kN, and 45.9 kN, respectively. The increase in loading capacity

compared to the un-repaired beam (SB-C) is up to 37.7%. The

flexural resistance of specimen SB-C-F1-IL has almost been

improved to equal to the reference beam SB-R. In comparison with

the simulated sea-sand beams repaired by sole SS, it is found that

the ICCP-SS technique showed its superior advantage (with up to

6.7% increase regarding to the flexural capacity). The reason for

this is because the ICCP-SS technique not only impedes further

corrosion of re-bars, but also recovers the strength loss of the

corroded specimens; while for the specimens repaired by SS

technique, corrosion of re-bars continues. Furthermore, it can be

found from comparison with beams repaired by sole ICCP or SS that

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the ICCP-SS technique has shown a superior advantage. However,

the different effects of small and large current densities are not

clear from results. More importantly, the beams repaired by ICCP-SS

technique showed similar loading capacities as the uncorroded

beams (SB and SB-R). The test results proved the effectiveness of

the ICCP-SS technique. However, the differences between the

specimens repaired by ICCP, SS and ICCP-SS techniques are not

sufficiently distinct. The reasons might be largely related to a short

ICCP operation duration and an insufficient amount of carbon fabric

mesh. Both the accelerated corrosion and the ICCP are only

operated for 130 days, which are rather short compared to the

common service life of RC structures (i.e. 50-100 years). The ICCP-

SS technique is a retrofitting method to ensure the durability of RC

structures, therefore its superiority will become more evident over a

longer period. In the next series of tests, this issue will be fully

considered to improve the experimental design.

Steel re-bars were taken out for mass loss measurement after bending tests

according to ASTM G1-03 [38]. The mass loss of steel-rebars were recorded. The

accelerated corrosion decreased the diameter of steel rebars, which was found to be

approximately 4.8% by loss measurement after tests. The original nominal area of

steel rebars was 157 mm2, and it was reduced to 142.3 mm2 after corrosion. Therefore,

the cross-section area of re-bars from corroded specimens without ICCP treatment

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(i.e. SB-C and SB-C-F1) is taken as 142.3 mm2 for capacity prediction due to the 260

days accelerated corrosion process. As for specimens containing chloride ions without

ICCP protection (i.e. SB-C-IS, SB-C-IS-R, SB-C-IL, SBC-IL-R, SB-C-F1-IS, SB-C-

F1-IS-R and SB-C-F1-IL), the rebar cross-section area is estimated as 149.6 mm2 (i.e.

assuming half mass loss at the end of 130 day comparing to corroded specimen

without ICCP treatment) since these specimens have also experienced 130 day

accelerated corrosion process.

4 Discussions on the cracking load and yielding loads

The cracking load (Pcr) is the load causing the initial crack of concrete. In the light of

ACI 318-14 [39], the equivalent section stiffness is derived by converting the elastic

modulus of steel re-bars (Es) to the elastic modulus of concrete (Ec). Similarly, for the

FRCM-strengthened beams, the effect of FRCM could be considered by converting

the elastic modulus of carbon fabric mesh (Ef) to the elastic modulus of concrete (Ec).

The ratio between the elastic modulus of re-bars and concrete (ns), and the ratio

between the elastic modulus of carbon fabric mesh and concrete (nf) as well as the

ratio between the elastic modulus of cementitious matrix and concrete (nm) are

determined by Eqs. (2-4). Afterwards, the neutral axis depth of the cracked section

(x0) is given by Eq. (5), and the cracking moment of inertia (I0) is derived by Eq. (6).

Finally, the cracking moment of FRCM strengthened beams (Mcr) can be obtained by

Eq. (7).

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(2)

(3)

(4)

(5)

(6)

(7)

where,

Af = area of FRP external reinforcement;

As = area of steel reinforcement;

Am = area of cementitious matrix;

b = width of the beam;

, modulus of rupture of concrete;

= specified compressive strength of concrete in cylinder, ;

= specified compressive strength of concrete in cube;

h = height of the beam;

h0 = distance from extreme compression fibre to centroid of steel tension

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reinforcement;

hf = distance from extreme compression fibre to centroid of carbon fibre tension

reinforcement;

hf = distance from extreme compression fibre to centroid of cementitious matrix

tension reinforcement;

= distance from centroidal axis of gross section, neglecting reinforcement, to

tension face, ;

= modification factor reflecting the reduced mechanical properties of lightweight

concrete, for normal weight concrete .

The predicted cracking loads are found to be 8.09 kN for reference beams (SB)

and 8.04 for SB-C, and 11.08 kN for FRCM-strengthened beams (SB-C-F1). The

cracking loads obtained from four-point bending tests were 10.9 kN of specimen SB-

C, and 14.5 kN for specimen SB. The cracking loads were all up to 19.1 kN in the

FRCM-strengthened beams. Compared the predicted and tested results, the increasing

ratio on the cracking loads is underestimated for the FRCM-strengthened beams.

The yielding load (Py) is the load when tensile steel re-bars reach the yield

strength. The yielding strain of steel re-bars (s) is 0.0019, obtained from steel tensile

tests according to ASTM E8/E8M [33]. Based on the basic bending theory, the strain

relations among concrete, steel bars, and carbon mesh are described in Eq. (8).

Meanwhile, the internal force equilibrium of the section is shown in Eq. (9). Based on

Eqs. (8) and (9), the neutral axis depth at steel yielding state (cy), the concrete strain

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(c) and the carbon fabric strain (fe) can be derived. Afterwards, the yield moment of

FRCM strengthened beams (My) is obtained according to Eq. (10).

(8)

(9)

(10)

where,

, compression force provided by concrete;

, modulus of elasticity of concrete;

, flexural moment provided by steel bars;

, flexural moment provided by carbon fabric mesh;

, tensile force provided by steel bars;

, tensile force provided by carbon fabric mesh;

Finally, the yielding load for reference beams (SB), corrosive beams (SB-C) and

FRCM-strengthened beams（SB-C-F1）were predicted to be 30.5 kN, 27.8 kN and

28.6 kN, respectively (see Table 5). The yielding loads obtained from four-point

bending tests were 41.9 kN for specimens SB, 27.3 kN for specimen SB-C, and 38.7

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397

398

399

400

401

402

403

404

405

406

407

408

409

kN for FRCM-strengthened beams (SB-C-F1) (see Table 5). Comparing the loading

capacities of SB and SB-C, capacity decrease due to corrosion is 34.8% from tests and

8.91% from calculation. Similarly, in comparison between specimens SB-C and SB-

C-F1, capacity increase due to C-FRCM strengthening is 41.7% from tests and 2.9%

from calculation, while capacity increase due to cathodic protection is 42.1% - 52.8%

from tests and 4.9% from calculation by comparing specimen SB-C, SB-C-IS, SB-C-

IS-R, SB-C-IL and SB-C-IL-R. There are three possible reasons causing the

differences between experimental and predicted results: (1) the corrosion of the re-

bars are not uniform; (2) the re-bars for weight measurement were taken from shear

span instead of constant moment span; (3) the material properties of C-FRCM

determined from tensile tests are different from its loading behaviour in bending tests

due to the different loading configuration and boundary conditions. However, since

stress-strain relationship before the slippage of fibre net may not be affected much,

the effect of this attribution might be limited.

5 Flexural capacity predictions

In this section, the flexural capacities from tests (Pexp) are compared with the design

flexural moment strengths given predicted by using ACI 440.2R-08 [26] (PACI 440) and

ACI 549.4R-13 [32] (PACI 549), as shown in Table 6. The design bending moment

capacity (Mn) is the combination of the flexural strength provided by the steel re-bars

(Mns) and the externally bonded C-FRCM (Mnf), as given in Eq. (11).

(11)

410

411

412

413

414

415

416

417

418

419

420

421

422

423

424

425

426

427

428

429

430

431

where,

, compression force provided by concrete;

, flexural moment provided by steel bars;

, flexural moment provided by carbon fabric mesh;

, compressive strain of unconfined concrete corresponding to

The key step is to assess the contribution of carbon fabric mesh at the ultimate

limit state. The measured specimen dimension and material properties were used in

the predictions, and all safety factors were set to be unity.

5.1 ACI 440.2R-08 [26]

The ACI 440 [26] is a design code for FRP – epoxy resin strengthened structures.

According to the design rules in ACI 440.2R-08, the effective design strain of FRP (

432

433

434

435

436

437

438

439

440

441

442

443

444

445

446

) at concrete crushing ( = 0.003) is calculated by Eq. (12), which is found to be

0.012. The debonding strain of externally bonded FRP ( ) is given by Eq. (13),

where Ef is the tensile modulus of elasticity of carbon fabric mesh equal to 223 GPa;

the debonding strain was found to be 0.012. According to the design criteria, if

, the specimen was failed by FRP debonding; otherwise, it is failed by

concrete crushing. According to calculated values of fe and fd, the failure mode for

the considered specimens is predicted to be FRP debonding. The calculated flexural

capacity (PACI440) of FRCM-strengthened beams (SB-C-F1) is 36.9 kN with debonding

failure, as presented in Table 6.

(12)

(13)

5.2 ACI 549.4R-13 [32]

The ACI 549 [32] is a design code specifically for the FRCM strengthening structural

design. The design procedure specified in ACI 549.4R-13 [32] is similar to the

guidance in ACI 440.2R-08. The only difference is that the effective tensile strain of

C-FRCM composites should be used in the prediction, instead of the CFRP properties.

As described in section 2.2, the material properties of the C-FRCM composites (see

447

448

449

450

451

452

453

454

455

456

457

458

459

460

461

462

463

Table 4 and Fig. 3) have been determined by the tensile coupon tests. The design

strain of C-FRCM composite (εfd) are defined as the average value minus one standard

deviation based on the test results, which is found to be 0.0104 according to the

tensile coupon test results presented in Session 2. The effective strain of C-FRCM at

failure (εfe) equals to the design strain (εfd), but smaller than 0.012, as given by Eq.

(14). Consequently, the effective tensile stress (ffe) of C-FRCM can be calculated in

accordance to Eq. (15). Therefore, the effective tensile strain (εfe) was determined to

be 0.0104, and the effective tensile stress (ffe) was found to be 2028 MPa based on Eq.

(15), where the tensile modulus of elasticity of C-FRCM equals to 195 GPa as

discussed in Session 2.2. Finally, the flexural capacity (PACI549) of FRCM-strengthened

beams (SB-C-F1) was calculated to be 33.5 kN with the predicted failure mode as

failure of C-FRCM composite. Please be noted that symbols εfe and ffe represent

effective strain and stress of CFRP in Session 5.1, but mean effective strain and stress

of C-FRCM composites in section 5.2.

(14)

(15)

5.3 Discussions on the theoretical and experimental results

The experimental flexural strength of SB-C-F1 specimen was 43.2 kN, and of FRCM-

strengthened beams with ICCP treatment (SB-C-F1-IS, SB-C-F1-IS-R, and SB-C-F1-

IL) ranged from 41.4 – 45.9 kN. The failure mode is slippage and fracture of fibre

mesh within the cementitious matrix and steel yielding followed by concrete crushing;

no FRCM debonding was observed. The design capacities for FRCM strengthened

464

465

466

467

468

469

470

471

472

473

474

475

476

477

478

479

480

481

482

483

484

485

beams were found to be 36.9 kN and 33.5 kN according to ACI 440.2R-08 [26] and

ACI549.4R-13 [32], respectively. Both design codes have found to underestimate the

flexural capacities of the beams tested in this study. The predicted failure mode of

ACI 440 is debonding failure of carbon fabric mesh from concrete substrate, and that

of ACI 549 is failure of C-FRCM composite. The underestimation of ACI 440 is

because the design rules in ACI 440 was developed based FRP-epoxy resin but not C-

FRCM. As for ACI 549, the inaccurate prediction is due to the different predicted

failure mode, as the design code cannot predict the combined failure of slippage and

fracture of carbon fibre mesh.

The effective tensile strain of carbon fibre mesh determined according to ACI

440.2R-08 is 0.012, while the effective tensile strain of C-FRCM composite

determined based on ACI 549.4R-13 is 0.0104. The measured strain at the ultimate

load was 0.007. However, it is difficult to accurately measure strains on the surface of

cementitious materials and carbon fibre meshes using conventional strain gauges [40].

The effective strain measured in the tests might be smaller the true value at the exact

location of fibre fracture.

Alternatively, it might be more straightforward to investigate the strengthening

capacity by using the effective stress instead of effective strain. It is found that using

effective stress could better predict the flexural capacities. According to ACI 440, the

effective tensile stress (ffe) of carbon fiber mesh is calculated to be 2754 MPa based on

the effective strain (εfe = 0.012) specified in ACI 440 and the measured elastic

modulus (Ef = 223 GPa). According to ACI 549, the effective tensile stress of the C-

486

487

488

489

490

491

492

493

494

495

496

497

498

499

500

501

502

503

504

505

506

507

FRCM composite ffe = 2028 MPa was obtained by multiplying the elastic modulus (Ef

= 195 GPa) and the effective strain (εfe = 0.0104). It should be noted that the value of

elastic modulus (Ef = 195 GPa) was based on the second part of the stress-strain curve

(see Fig. 3), as defined by ACI 549 [32]. This value is only 65% of elastic modulus of

the first part of the stress strain curve (i.e. 299 GPa). This could lead to rather

conservative prediction for the effective stress even though the effective strain is the

mean value obtained from material tests. From the tensile test results (see Fig. 3), it

can be seen that the ultimate stress of C-FRCM is 2028 MPa. Based on the

experimental loading capacity of the beam and measured strain of steel re-bar, the

strength developed in C-FRCM composite was derived to be 1447 MPa when the

beam reaches ultimate loads. This is close to 70% of the ultimate stress of C-FRCM

measured from tensile tests. According to the measured reading of strain gauges on

the steel rebar surface, re-bars have reached the strain hardening region at ultimate of

bending tests. If the resistance provided by the C-FRCM and steel-rebars are

calculated based on the measured strain, the capacity of the ICCP-SS beam is found to

be 44.5 kN, which is rather close to the test results (41.4 – 45.9 kN).

Based on the flexural theory, the reduction of ultimate loading capacity due to the

loss of re-bar section is calculated to be 8.8% (i.e. comparing SB and SB-C); the

reduction of loading capacity is 29.2% - 33.7% based on test results. The reduction of

experimental yielding loads due to corrosion is 34.8% as mentioned in Section 4,

which is rather consistent with the reduction of experimental ultimate loads. The main

reasons causing the difference are the non-uniform corrosion of the re-bars. Please

508

509

510

511

512

513

514

515

516

517

518

519

520

521

522

523

524

525

526

527

528

529

also note that the re-bars for weight measurement were taken from shear span instead

of constant moment span. In comparison of specimens repaired with and without C-

FRCM (i.e. SB-C and SB-C-F1), the increase on the ultimate loads is 29.3% from

tests and 15.7% from ACI 440 prediction. Similarly, comparing specimens SB-C, SB-

C-IS, SB-C-IS-R, SB-C-IL, SB-C-IL-R, capacity increase due to cathodic protection

is 26.6% - 40.7% from tests and 13.2% from calculation. Please be noted that the

prediction capacity increase due to ICCP was calculated using the difference in

measured weight loss of corrode re-bars. It is found that predicted capacity increase

due to ICCP is less than the experimental results. It might be attributed to the more

serious pitting corrosion in the constant moment region of the corroded specimens

without ICCP treatment (specimen SB-C) which were not identified in the mass loss

measurement.

6 Conclusions

A dual-functional retrofitting method is investigated for sea-sand reinforced concrete

beams. This method combines the merits of impressed current cathodic protection

(ICCP) and structural strengthening (SS) techniques. This paper presents the

experimental program and discusses the test results. The experimental program

includes an accelerated corrosion procedure, the ICCP process, and the bending tests.

From the test results, it is found that the C-FRCM is capable of being exposed to high

current densities up to 80 mA/m2 without mechanical bonding. The C-FRCM

composite can be used to strengthen sea-sand RC beams,

530

531

532

533

534

535

536

537

538

539

540

541

542

543

544

545

546

547

548

549

550

551

maintaining the structural integrity and increasing the ultimate

strength of damaged beams. The ultimate strengths of reinforced

concrete beams repaired by ICCP, SS and ICCP-SS techniques are,

respectively, approximately 29.3%, 40.7% and 37.7% greater than

corroded beams without any repairing, the decrease in strengths

compared to the uncorroded reference beams by the ICCP, SS and

ICCP-SS techniques are 12.5%, 11.5% and 15.2%, respectively. The

ICCP-SS technique works effectively, but its superior merit has not

been fully explored in this study; this might be largely attributed to

the insufficient corrosion level and the limited ICCP period. The

crack loads, yielding loads and ultimate loads of specimens from

tests were compared with those obtained from theory calculations.

The effects on loading capacities due to re-bar corrosion, C-FRCM

strengthening and cathodic protections were discussed based on

both experimental and predicted results. The reduction of

experimental yielding loads due to corrosion is rather consistent

with the reduction of experimental ultimate loads, but greater than

the predictions. The main reasons causing the difference are the non-uniform

corrosion of the re-bars and that the re-bars for weight measurement were taken from

shear span instead of constant moment span. In future, more efforts are

needed to optimize the applied current densities and the amount of

strengthening material. Longer operation periods for the corrosion

552

553

554

555

556

557

558

559

560

561

562

563

564

565

566

567

568

569

570

571

572

573

and ICCP process, as well as the effects of different current densities

in ICCP should be considered to determine the durability

performance of sea-sand RC structures.

Acknowledgements

We would like to thank the support from the Chinese National Natural Science

Foundation (51778370, 51538007), Natural Science Foundation of Guangdong

(2017B030311004), the Key Project of Department of Education of Guangdong

Province (No.2014KZDXM051), the Shenzhen science and technology project

(JCYJ20170818094820689).

NotationsAf = area of FRP external reinforcement;

As = area of steel reinforcement;

b = width of the beam;

= neutral axis depth at steel yielding state;

, compression force provided by concrete;

, modulus of elasticity of concrete;

Es = elastic modulus of steel re-bars；Ef = elastic modulus of carbon fabric mesh;

Efrcm = tensile modulus of elasticity of cracked C-FRCM;

574

575

576

577

578

579

580

581

582

583

584

585586

587

588

589

590

591

592

593

594

= specified compressive strength of concrete in cylinder, ;

= specified compressive strength of concrete in cube;

ffe = effective tensile stress of C-FRCM;

ffu = ultimate tensile strength of C-FRCM;

, modulus of rupture of concrete;

h = height of the beam;

h0 = distance from extreme compression fiber to centroid of steel tension

reinforcement;

hf = distance from extreme compression fiber to centroid of carbon fiber tension

reinforcement;

I0 = cracking moment of inertia;

Mcr = cracking moment of FRCM strengthened beams;

My = yield moment of FRCM strengthened beams;

Mn = design bending moment capacity;

Mns = flexural strength provided by the steel re-bars;

Mnf = flexural strength provided by the externally bonded C-FRCM;

, flexural moment provided by steel bars;

, flexural moment provided by carbon fabric mesh;

n = number of layers of carbon fabric mesh;

595

596

597

598

599

600

601

602

603

604

605

606

607

608

609

610

611

612

613

, ratio between the elastic modulus of re-bars and concrete;

, ratio between the elastic modulus of carbon fabric mesh and concrete;

Pcr = cracking load；Py = yielding load;

PACI440 = flexural capacity of FRCM-strengthened beams in light of ACI 440.2R-08;

PACI549 = flexural capacity of FRCM-strengthened beams in light of ACI 549.4R-13

Pu = flexural capacity obtained from tests

, tensile force provided by steel bars;

tf = thickness of carbon fabric mesh;

, tensile force provided by carbon fabric mesh;

x0 = neutral axis depth of the cracked section;

= distance from centroidal axis of gross section, neglecting reinforcement, to

tension face, ;

; factor relating strength of equivalent rectangular compressive

stress block to specified compressive strength of concrete;

, factor relating depth of equivalent rectangular compressive stress

614

615

616

617

618

619

620

621

622

623

624

625

626

627

628

629

block to depth of neutral axis;

c = concrete strain;

, compressive strain of unconfined concrete corresponding to ;

εcu = 0.003, ultimate compressive strain of concrete;

εfe = effective tensile strain of carbon fabric mesh;

εfd = design tensile strain of carbon fabric mesh;

= tensile strain of C-FRCM at a stress level equal to 0.9ffu;

= tensile strain of C-FRCM at a stress level equal to 0.6ffu;

s = yielding strain of steel re-bars;

= modification factor reflecting the reduced mechanical properties of lightweight

concrete, for normal weight concrete .

630

631

632

633

634

635

636

637

638

639

640

641

642

Figures

Figure 1. Detailed dimensions of beam specimens (all dimensions in mm).

Figure 2. Failure mode of C-FRCM observed from uniaxial tensile test.

Figure 3. Stress – strain response curve of C-FRCM obtained from uniaxial tensile

test.

Figure 4. (a) Schematic illustration of impressed current cathodic protection (ICCP)

of reinforced concrete beams.

(b) Beam with ICCP set-up.

Figure 5. Configuration of simply supported beam tests (all dimensions in mm).

Figure 6. Potential of re-bars during operation of ICCP.

Figure 7. Typical failure mode of the tested beams (specimen SB-C-F1-IL).

Figure 8. Responses of the loading capacity-deflection.

643

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646

647

648

649

650

651

652

653

654

655

656

657

658

659

Figure 1.

Figure 2.

0

1000

2000

3000

4000

0 0.005 0.01 0.015 0.02

Stre

ss (M

Pa)

Strain

Carbon fabricC-FRCM

Figure 3.

660

661662663664

665

666667668669

670671672673674675

(a)

(b)Figure 4.

Figure 5.

676

677

678679

680

681682683684685

686687688

Figure 6.

Figure 7.

689690

691692693694695

696

697698699700

0

10

20

30

40

50

60

0 5 10 15 20 25

Load

s, P

(kN

)

Deflection, δ (mm)

SBSB-R

SB-C-F1

SB-C

(a) Effect of strengthening of C-FRCM composite

0

10

20

30

40

50

60

0 5 10 15 20 25 30 35

Load

s, P

(kN

)

Deflection, δ (mm)

SB SB-R

SB-C

SB-C-IL-R

SB-C-ISSB-C-IS-R

SB-C-IL

(b) Effect of ICCP using C-FRCM anode

0

10

20

30

40

50

60

0 5 10 15 20 25 30 35

Load

s, P

(kN

)

Deflection, δ (mm)

SB SB-RSB-C-F1-IL

SB-C-F1-IS

SB-CSB-C-F1-IS-R

(c) Effect of ICC-SS using dual functional C-FRCM compositeFigure 8.

701702703704

705706707

708709710711

Tables

Table 1. List of beam specimens.

Table 2. Ingredients of concrete mixture.

Table 3. Ingredients for the cementitious material.

Table 4. Material properties of concrete, steel re-bars, carbon fabric mesh, and

cementitious matrix.

Table 5. Yielding loads obtained from tests and prediction

Table 6. Test results and the comparison with ACI 440 and ACI 549.

712713

714

715

716

717

718

719

720

721

722

723

Table 1.

Noted: SB = simple supported beams; C = chloride; F1= externally bonded C-FRCM

with one lay of carbon fabric mesh; IS = applied small current density in the ICCP

process; IL = applied large current density in the ICCP process.

Table 2.

Cement(kg)

Fine aggregate

(kg)

Coarseaggregate

(kg)

Water(kg)

Superplasticizer

(ml)

NaCl(by the weight of cement, %)

1 1.29 2.88 0.39 0.01 3

Table 3.

Composition By the weight of cement

SpecimensNaCl

(by the weight of cement, %)

Parameters in each technique

Layers of carbon fabric mesh in the SS

Current densities in the ICCP (mA/m2)

SB 0 0 0SB-R 0 0 0SB-C 3 0 0SB-C-F1 3 1 0SB-C-IS 3 0 26SB-C-IS-R 3 0 26SB-C-IL 3 0 80SB-C-IL-R 3 0 80SB-C-F1-IS 3 1 26SB-C-F1-IS-R 3 1 26SB-C-F1-IL 3 1 80

724

725

726

727

728

729

730

731

732

733

734

735

736

737

(%)

Cement 100.00Silica fume 22.22

Re-dispersed polymer 11.11Water 50.00

Carboxymethylcellulose 0.25Defoamer 0.53

Superplasticizer 1.20Chopped carbon fibre 1.00

Table 4.

MaterialThickness/ Diameter

(mm)

Yield stress (MPa)

Tensile strength(MPa)

Compression strength(MPa)

Young’s Modulus

(GPa)

Tensile strain(%)

Concrete --- ---- --- 53.0^ --- ---

Re-bars(HRB400)

10 382 544.3 --- 200 ---

Carbon fabric mesh

0.207 --- 3519 --- 223 1.58

Cementitious matrix

--- --- --- 37.9 76 ---

C-FRCM --- --- 1584 --- 195 (299) 1.04

Note: ^ Test results were from concrete cube tests in 150 mm×150 mm.

( ) Value in the blanket is the elastic modulus of the first part of stress-strain curve

Table 5

SpecimensYielding loads from tests

(kN)Yielding loads from prediction

(kN)

SB 41.9 30.5SB-R 41.2 30.5

738

739

740

741

742

743

744

745

746

747

748

SB-C 27.3 27.8SB-C-F1 38.7 28.6SB-C-IS 40.1 29.2SB-C-IS-R 42.0 29.2SB-C-IL 42.0 29.2SB-C-IL-R 39.1 29.2SB-C-F1-IS 39.4 30.0SB-C-F1-IS-R 38.0 30.0SB-C-F1-IL 44.1 30.0

Table 6

Specimens

Experimental

ultimate

loads

(kN)

Decrease in strength

compared to control

beams (SB and SB-R)

(average load = 48.8

kN)

(%)

Increase in

strength compared

to control beam

(SB-C)

(load = 33.4 kN)

(%)

SB 50.4 --- 50.9 32.2 32.2

SB-R 47.2 --- 41.3 32.2 32.2

SB-C 33.4 31.5 --- 29.4 29.4

SB-C-F1 43.2 11.5 29.3 36.9 33.5

SB-C-IS 42.3 13.3 26.6 30.8 30.8

SB-C-IS-R 46.9 3.9 40.4 30.8 30.8

SB-C-IL 47.0 3.7 39.9 30.8 30.8

SB-C-IL-R 43.2 10.1 29.3 30.8 30.8

SB-C-F1-IS 41.4 11.5 24 39.3 34.9

SB-C-F1-IS-R

44.0 9.8 31.7 39.3 34.9

SB-C-F1-IL 45.9 5.7 37.7 39.3 34.9

749750

751

752

753

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