STRENGTHENING OF BRICK MASONRY ARCHES WITH EXTERNALLY BONDED STEEL REINFORCED COMPOSITES

Antonio Borri1, Paolo Casadei2, Giulio Castori3, John Hammond4

ABSTRACT

The objective of this study is to investigate the efficiency of an innovative technique for strengthening masonry

arches, based on the use of high strength steel cords embedded in either an epoxy (SRP: Steel Reinforced

Polymer) or mortar matrix (SRG: Steel Reinforced Grout). Ten prototypes of brickwork arches strengthened by

composite laminates were tested under a monotonic vertical load applied at the quarter span. Load tests were

performed to compare the behavior up to collapse of strengthened masonry arches; the influence of the types of

reinforcement (steel and carbon fibers) and matrices (epoxy and cementitious), as well as location of the

strengthening layer (intrados, extrados and both) and the presence of anchorage systems has been investigated.

The experimental results highlight the enhanced strength of the arches reinforced with steel cords, as well as the

role of the mechanical anchoring with regard to the resulting final strength.

CE Database subject headings: Arches; Masonry; Composite materials; Steel cords.

INTRODUCTION

Among the various structural components of monumental buildings, masonry arches merit particular attention.

They are very widespread in Italian historical centers, and their preservation as part of the cultural heritage is a

very timely subject. Because of the degradation owing to time and other accidental causes (such as earthquakes),

these structures suffer several types of damage. These necessitate the application of strengthening materials and

repair techniques designed to re-establish initial structural performances and prevent the brittle collapse of the

masonry due to possible future hazardous conditions. Traditional retrofit methods (Day 1995; Boughton et al.

1997) used to prevent premature or sudden collapse include: single- or double-sided jacketing with cast-in-situ

1 Professor, Department of Civil and Environmental Engineering, University of Perugia, via Duranti 93, 06125

Perugia, Italy.

2 Ph.D., PE, Research-Design Engineer, TECINN S.r.l., Perugia-Milano, Italy

3 Ph.D., Researcher, Department of Civil and Environmental Engineering, University of Perugia, via Duranti 93,

06125 Perugia, Italy. Tel: +390755853906, Fax: +390755853897, E-mail: gcastori@strutture.unipg.it.

4 Vice President, Hardwire LLC, Pocomoke City, Maryland, USA.

reinforced concrete or shotcrete, in combination with steel reinforcement (e.g., in the form of two-directional

welded mesh); the use of reinforced grouted injections or crack stitching ties; internal or external post-

tensioning with steel ties to link structural elements together into an integrated three-dimensional system. All

above techniques may guarantee an adequate increment in strength and stiffness, but are often short-lived, labor-

intensive, and usually violate the aesthetic, conservation and restoration requirements. Such problems have led

researchers (Foraboschi 2001; Valluzzi et al. 2001; Borri and Castori 2004) to investigate new retrofit solutions

using innovative materials such as fiber-reinforced polymer (FRP) composites in the form of bonded surface

reinforcements. FRP reinforcement offers the designer an outstanding combination of properties, including low

self-weight, long-term durability, and high strength. Moreover, the use of these materials does not alter the

natural behavior of the structure since they do not add mass. In addition, they are removable, and they can be

made either invisible or visible, in order to comply with modern restoration requirements. On the other hand,

their low fire resistance and their relatively high cost may represent obstacles for their widespread use.

A very promising development of interest in the use of composites regards a new family of innovative

materials (Hardwire LLC 2002), based on ultra high strength steel filaments twisted to form cords and

embedded in either an epoxy (referred to as Steel Reinforced Polymer, SRP) or mortar matrix (referred to as

Steel Reinforced Grout, SRG). The use of these materials in seismic upgrading presents several interesting

aspects which make them very attractive for retrofit applications. These can be summarized as follows: a) steel

cords have an inherent shear strength that can simplify problems concerning connections and anchorages; and b)

impregnation with mortar matrix - instead of epoxy matrix - may overcome fire endurance and physical and

chemical compatibility problems (e.g., reduction of masonry permeability). On the other hand, SRP and SRG

are fairly new technologies and, at present, few studies are available on the use of such materials as

strengthening systems (Wobbe et al. 2001; Casadei et al. 2005; Huang et al. 2005; Matana et al. 2005; Prota et

al. 2006; Borri et al. 2007; Cancelli et al. 2007). The intention of this study is to compare the behavior up to

collapse of strengthened masonry arches through the study of types of reinforcement (steel and carbon fibers)

and matrices (epoxy and cementitious), location of the strengthening layer (intrados, extrados and both) and the

presence of anchorage systems.

BEHAVIOR OF THE STRENGTHENED ARCHES

The stability as well as the safety of curved structural members under a given loading condition are strongly

dependent on the geometry of the structures and on the mechanical characteristics of the constituent material.

Masonry has a well-known negligible tensile strength, so the safety condition for masonry arches (or vaults) is

achieved when the line of thrust, coincident with the funicular polygon, is kept within each section of the arch

itself. When the resultant of the internal forces moves outside the central core, the section cracks, signaling the

onset of a phase of high deformations. This results in the formation of a plastic hinge, which exhibits the

crushing of a limited portion of the masonry at the compressed edge of the arch. If sufficient hinges form to turn

the arch into a mechanism, then collapse is possible.

The use of composite materials can modify the failure mode of masonry arches and significantly increase

their load-carrying capacity. Reinforcements externally bonded to the masonry surface enable masonry

structures to carry substantial tensile stresses, eliminating their greatest mechanical shortcoming at an acceptable

cost. Therefore, it is possible to avoid the brittle collapse mechanism of such structures, typically caused

(assuming that the abutments are rigidly fixed) by the formation of four hinges. Depending on the position of the

laminate, in fact, the formation of the forth hinge can be prevented. More specifically, in the case of extrados

(outer surface) strengthening the line of thrust can fall outside the lower edge of the arch (Figure 1a) without any

subsequent structural collapse. As a result, in the case of a vertical load applied at a quarter span, haunch hinge

formation is prevented and the arch becomes a statically determinate structure (it is a three-hinge arch)

consisting of two curved beams strengthened on their upper sides. Conversely, in the case of a structure

strengthened at the intrados (inner surface), the thrust line can fall outside the upper edge of the structure (Figure

1b) and the reinforcement prevents the formation of the forth hinge close to the load point.

Consequently, in both cases, collapse is due to other mechanisms, which are dependent on the limits of

strength of the constituent materials (masonry and reinforcement) and on their structural interactions at the local

level (i.e. bond and localized shear). Depending on the position and amount of the reinforcement, the only

remaining types of structural failure may be due to: masonry crushing, reinforcement rupture or debonding, and

shear sliding along a mortar joint. Analytical formulations available in literature for evaluating the load-carrying

capacity of masonry arches strengthened with composites are herein reported. All of them assumed a linear

elastic behavior for the reinforcement and a rectangular stress-block law - whose depth (x) is one third of the

height of the section (t) - for the masonry. Even though the latter assumption is not properly introduced when

dealing with masonry structures reinforced with composites, the deviations between the experimental and

assumed values of x are negligible (Foraboschi 2004).

The two mechanisms of masonry crushing and reinforcement rupture, have been studied and formulated by

Triantafillou (1998), who defined the moment capacity of strengthened sections under combined compressive

and bending stresses, similarly to the one assumed for RC structures. According to such formulation, the

formulas governing the behaviour are given by (Triantafillou 1998):

2518

= ⋅ ⋅Rd , mc M , uM f l t (1)

256

= ⋅ ⋅ ⋅Rd , rr reinf,uM f l tρ (2)

where MRd,mc = moment capacity for masonry crushing; fM,u = masonry crushing stress; l = width of the cross

section; MRd,rr = moment capacity for reinforcement rupture; freinf,u = reinforcement rupture stress; and ρ =

reinforcement area fraction. The other two mechanisms, namely shear sliding and debonding of the

reinforcement, are based on the local interaction between the constituent materials, so that, the relative design

parameters, can be obtained from the mechanical properties of each single constituent material. For the shear

sliding mechanism, it is supposed that the corresponding resisting strength (VRd) is given by (Foraboschi 2004):

65

⋅ ≤⎧⎪

= ⎨ ⎛ ⎞+ ⋅ >⎪ ⎜ ⎟⎝ ⎠⎩

Rd

N if M/N t/2V MN if M/N t/2

t

μ

μ (3)

where μ = masonry friction coefficient; N = design axial load; and M = design bending moment.

As a result, failure occurs once the shear action (V) reaches the sliding resistance (VRd) in a generic cross

section. As for debonding, it is supposed that the failure is correlated to the presence of a component

perpendicular (σ) to the directrix of the curve, whose value is given by (Borri and Castori 2004):

65

reinf

M Nt

b Rσ

−=

⋅ (4)

where breinf = width of the reinforcement; and R = radius of curvature.

The crucial point is to establish the limit (σRd) of the σ. For this purpose, the theoretical analysis has fully

incorporated the experimental results. σRd can be measured by in situ tests (pull-off tests) or obtained from the

technical literature. Failure occurs once σ reaches σRd in a generic cross section.

EXPERIMENTAL PROGRAM

Characterization of the materials

The arches were constructed using concrete bricks (200x100x50 mm) and a hydraulic lime mortar (ratio

sand/binder = 5/2 in volume). The mechanical properties of the bricks were obtained by means of compression

and bending tests, each of which was carried out on six samples. Uniaxial compression tests gave a mean

strength of 43.3 MPa, whereas the mean value of the bending tensile strength was 10.9 MPa. As for the

hydraulic lime mortar, compression tests carried out on mortar prisms after 28 days of curing gave a mean

strength equal to 0.72 MPa. The composite system used to strengthen the arches included carbon fibers and steel

cords (Table 1). The steel cords have a brass coating, which enhances the corrosion resistance, and are arranged

parallel to one another and held in place by knit yarns forming an appropriate fabric pattern (Hardwire LLC

2002). The yarns control the spacing of the cords. Accordingly, two steel tapes of different densities and

morphology were used for this project. The first was a low-density cord tape (Type 1), used for intrados

strengthening, containing 4 cords per in. (0.16 cords per mm), with each cord formed by three straight wire

filaments wrapped by another two filaments at a high twist angle. The second was a medium-density cord tape

(Type 2), used for extrados strengthening, containing 12 cords per in. (0.47 cords per mm), with each cord

formed by three straight wire filaments wrapped by another filament at lower twist angle.

As for carbon fibers, all the data were provided by the manufacturers (Toray Industries 2003). Conversely,

coupon specimens were tested in tension to determine the mechanical properties of steel cords and to draw their

stress-strain curves. Test results indicated that the material behaves linearly to failure and there is practically no

yielding of the steel. A two-component epoxy resin and a high performance cementitious grout were used to

impregnate and bond the steel cords and carbon fibers. The mechanical properties of the matrices, acquired from

experimental tests, are shown in Table 2.

Interaction between masonry and steel cords

To investigate the mechanisms of local interaction among the constituent materials, adhesion tests on samples

strengthened by SRP and SRG were carried out for both loads parallel (direct shear tests) and loads

perpendicular (pull-off tests) to the bond surface (Figure 2). Direct shear tests revealed that while all SRG

specimens failed by debonding in the layer of grout, shearing within the bricks occurred with SRP reinforcement

(Figure 3a). The mean values of the shear strength were 1.88 and 1.21 MPa for SRP and SRG specimens,

respectively. Even in the case of the pull-off tests, all SRP specimens failed in the substrate, whereas SRG

specimens failed in the interface (Figure 3b). The mean values of the shear strength were 1.57 and 1.29 MPa for

SRP and SRG specimens, respectively. This indicated, according to fib provisions (Fib Bulletin 14), that both

for direct shear and pull-off tests, the failure of SRP specimens is a cohesion failure, while an adhesion failure

characterized the SRG specimens. Thus, while in the case of SRG strengthening the obtained values are

effectively equal to the bond strength between cords and masonry, in the case of SRP strengthening they

correspond to the masonry shear or tensile strength.

Tests on masonry arches

To study the behavior of the strengthened arches, ten specimens, built with concrete bricks arranged in a single

layer (100 mm of thickness) were tested under monotonic vertical loads applied at a quarter of their span. A

catenary curve (parabolic shape), typical of many already existing arches - as it is the ideal form to improve the

stability of the structure under loading - was chosen as directrix of the arches. A span of 1980 mm therefore

means that the height of the arch is 490 mm above the springing level. As shown in Figure 4, the supporting

structure was constructed of concrete abutments on a steel C-section base, fixed to the floor using threaded bars.

In order to measure the springing thrust, a bridge bearing and a reaction frame were used at either side. The

bridge bearing, on which the concrete abutment rests, provided the vertical reaction, allowing horizontal

superstructure motions. Horizontal reaction was supplied by a reaction frame, built with two angle iron brackets

(fixed to the floor), a steel beam and steel flat plates assembled together. A 100 kN load cell placed between

reaction frame and concrete abutment was used to measure the springing thrust. In addition, a total of eight

Linear Variable Displacement Transducers (LVDTs) were used to register deflections, while twelve strain

gauges were placed near the predicted hinge position to ensure the maximum strain was recorded.

As already noted, the arches were all tested by varying the position of reinforcement (extrados, intrados, both

intrados and extrados), type of matrix (epoxy and cementitious) and reinforcement (steel cords or carbon fibers),

number of plies (one or two) and boundary conditions. All the specimens, except for the control (UN.01), were

strengthened using a different combination of the above test variables (Table 3).

A single ply of laminate (150 mm wide) was applied to each arch, except in the case of arch IN.03 where two

plies of laminate were used. Steel cords were installed following the recommendations of ACI 440.2R-02 (ACI

440) provisions for FRP materials. The soffit or the extrados of the arches was first abrasive-blasted to ensure

proper bonding of the composite system. With the surface roughened and cleaned, the first layer of epoxy (or

grout) was applied directly, without primer and putty coatings (typically used in standard CFRP applications). A

rib-roller was then utilized to press on the tape to ensure epoxy (or grout) impregnation and encapsulation of

each cord. Finally, the excess resin (or grout) was spread with a putty-knife to create an even surface. For the

two ply application, once the first ply was in place and the excess resin (or grout) leveled, the second ply was

applied, following an identical procedure. Moreover, it is known that the steel cords have the potential to

address the shortcomings (low shear strength ability of FRP, possibilities of electrical incompatibility, stress

concentration, etc.) associated with the use of mechanical anchoring for FRP applications. Therefore, in two

cases (EX.03 and IN.03 tests) steel anchors were adopted, in addition to the reinforcement. In the first case, in

order to prevent slip, two angle plates were used to anchor the ply to the abutments (Figure 5a), while in the

second case, in order to delay premature peeling, flat plates, bolted to the bricks (every three bricks,

approximately 150 mm apart), were used to secure the ply to the arch soffit (Figure 5b).

TEST RESULTS

In the following, the results of the experimental tests carried out on the arches are grouped according to the type

of arrangement of the reinforcement on the surface of the masonry.

Unstrengthened arch

As was expected from the materials characterization, the unreinforced arch demonstrated brittle failure, due to

the formation of four hinges (arch displacement mechanism). The failure load, used as reference value for the

comparison with the strengthened arches, was equal to 0.7 kN.

Strengthening at Extrados

Due to a setup problem, arch EX.01 showed a notable lateral displacement of both abutments, which did not

allow any further increment in load (the ultimate load was 9.2 kN). Conversely, arches EX.02 and EX.03

exhibited the same failure mechanism (shear sliding), the difference being that the presence of steel anchors

(arch EX.03) allowed the sliding between brick and mortar in a different location. This resulted in the first joint

closest to the springer (Figure 6a) rather than the first joint closest to the edge of the steel anchor (Figure 6b); in

both cases such collapse occurred without any warning. The ultimate load was 13.6 and 23.4 kN for arches

EX.02 and EX.03, respectively. Finally, arch EX.04 showed the same failure mode (sliding in the first joint

closest to the springer) but a lower ultimate load capacity (11.5 kN). Due to the presence of undesirable

variables (such as handwork) that may have arisen from the construction of the specimens, this fact may be

attributed to a lower value of the masonry friction coefficient (μ).

Strengthening at Intrados

The arches strengthened at their intrados exhibited two different patterns of collapse: laminate rupture and

laminate debonding. More precisely, in specimens IN.01 and IN.02, employing a low-density tape able to ensure

a good interface bond at the expense of less tensile strength, failure was dictated by the rupture of the

reinforcement under the loaded section (Figure 7a); in both cases such collapse was brittle. The ultimate loads

were 16.2 and 14.7 kN for arches IN.01 and IN.02, respectively. Arch IN.03, reinforced with two plies of

laminate, failed due to laminate debonding. Detachment of the reinforcement started between two steel anchors

(Figure 7b) - probably initiated by the presence of a bonding defect - and then progressed towards the loaded

section. The presence of steel anchors delayed such a premature peeling and the complete detachment of the

laminate occurred only after a substantial increase of the load-carrying capacity of the arch (21.7 kN).

Finally, arch IN.04, reinforced with carbon fibers, exhibited the same failure mode (laminate debonding), but

a lower ultimate load capacity (12.3 kN). In this case, as well as in that of arch IN.04, the structure did not reach

a state of collapse, since the reinforcement contributed in holding the bricks together during the last phase.

Strengthening at Extrados and Intrados

Arch IN+EX.01 failed due to the sliding along the mortar joint close to the point of application of the load,

resulting in the consequent detachment of the reinforcement from the masonry (Figure 8). Under such

conditions, debonding progressed as long as the reinforcement contributed to holding the bricks together, then

the arch load-carrying capacity (33.0 kN) decreased suddenly and the specimen collapsed under its own weight.

ANALYSIS OF THE RESULTS

The experimental investigation carried out on the strengthened arches is aimed at assessing the potential of steel

cords to provide a strengthening system alternative to traditional techniques and to FRP laminates. It represents

a first step toward the development of a novel strengthening material system for structural upgrade. The analysis

of the test results is conducted first with respect to arches strengthened with equivalent reinforcement type (SRP

or SRG); and then, arches using the same strengthening arrangement with different materials (steel cords and

carbon fibers) are compared. Remarks on the influence of different matrix type (epoxy and cementitious) and

mechanical anchoring are also presented while experimental-theoretical comparison in terms of strength are

reported elsewhere (see Section below). According to such an analysis the following can be highlighted:

1. The position of the steel cords seems to have a nontrivial influence in the arch behavior, particularly when

comparing arches without mechanical anchors. In spite of lower values of the reinforcement ratio, the intrados

arrangement was more effective than extrados arrangement in terms of strength; the trend was inverted in terms

of ultimate deflections. A comparison (Figure 9 and Figure 10) between arches with equivalent reinforcement

type (due to a setup problem, arches IN.01 and EX.01 are not comparable) highlights that the arch strengthened

at the intrados (arch IN.02) provided an ultimate strength approximately 10% larger than that strengthened at the

extrados (arch EX.02) with deflections approximately 45% smaller;

2. Strength increases provided by steel cords were greater than those obtained using carbon fibers for both

intrados and extrados strengthening applications. Regardless of the type of matrix (epoxy or cementitious), steel

cords permitted the attainment of strength increases ranging between 18 and 32%, while increases of ultimate

deflections were approximately 18% and 250% larger for extrados and intrados applications, respectively.

3. As for the intrados strengthening, the cementitious grout seemed to be more effective than the epoxy resin

in engaging the masonry substrate. In final analysis, the SRG ultimate strength (arch IN.01) was approximately

10% more than that corresponding to the SRP (arch IN.02), while the maximum strain recorded in the SRP at

about 85% of the ultimate load (when gauges have started to produce unreliable data) was approximately 50%

lower than that corresponding to the SRG;

4. A significant role of the mechanical anchoring (arches EX.03 and IN.03) was observed. Compared with the

arches strengthened at the intrados (arches IN.01 and IN.02), arch IN.03 permitted the attainment of strength

increases ranging between 34 and 48%, while the ultimate strength of arch EX.03 was approximately 72%

larger than that corresponding to the arch EX.02;

5. Important information is also provided by the analysis of load cell readings at the springings of the arch.

By comparing the load-springing thrust curves (Figure 11) it can be observed that, in terms of a reduction in the

springing thrust, the sensitivity of the arch to the extrados reinforcement was much lower than that to the

intrados reinforcement. Yet, neither the intrados reinforcement alone nor the simultaneous intrados and extrados

reinforcement were sufficient to bring the springing thrust down to zero, i.e. to transform the arch into a beam.

COMPARISON BETWEEN THEORETICAL AND EXPERIMENTAL RESULTS

A comparison between experimental results and the predictions obtained by analytical formulations (Eqs. (1) to

(4)) was made to evaluate the possibility of the use of such formulations in predicting the behavior of the

strengthened arches. The analytical formulations presented in the paper could represent in fact a first step for the

development of code recommendations for the design of strengthening of masonry arches using steel cords.

This comparison reveals good agreement between the experimental data and theoretical predictions for the

collapse mode and, consequently, for the corresponding load-carrying capacity (Table 5). Specifically, arches

EX were predicted to fail due to sliding, as was also determined experimentally, and the maximum deviations

between the experimental and analytical load was found to be in almost all cases no more than 11%. A single

exception was in the case of arch EX.01, where the error of the model was 32%. However, as already

mentioned, this failure mode was due to a test setup problem not considered in the theoretical analysis, and

therefore such a disagreement can be disregarded. It should be noted that, due to the lack of an experimental

value, the maximum value of the sliding resistance (VRd) was calculated assuming μ = 0.5 (Heyman, 1982).

Conversely, with regard to series IN, while the analytical model properly simulated the experimental behavior of

the arches, which failed by laminate rupture (maximum deviations between the calculated and measured values

being no more than 16%), the agreement between experimental and theoretical results was less satisfactory for

arch IN.03, which failed by laminate debonding. This error may be explained by the fact that debonding in

masonry arches allows for the presence of a multiaxial stress state. This combines the normal stresses (σ), whose

values can be calculated analytically (see Eq. (4)), with the tangential stresses (τ), whose values have been

disregarded in the adopted model. The model should be refined so that it is able to assess the resultant state of

stress, acting in a generic cracked section, by considering the contribution both of normal and tangential stresses

(Borri and Castori 2004). Finally, good agreement between theoretical and experimental results was even

detected for arch IN+EX.01. The error of the model was in fact 6%. Conversely, as for the failure mode, the

model has shown to be effective in predicting the initial sliding mechanism, but it has not been able to predict

the consequent laminate debonding.

4 CONCLUSIONS

The following conclusions are deduced from the experimental results:

• SRP/SRG composite materials, similar to FRP in terms of ease of installation, allow the same applications,

reducing installation and material costs;

• In spite of their lower mechanical properties versus carbon fibers, steel cords performed better than CFRP in

terms of ultimate load capacity and ultimate deflection for both intrados and extrados strengthening

applications;

• Mechanical anchoring, which is usually impractical for FRP applications, was shown to successfully improve

the overall performance by allowing a substantial increase in terms of ultimate load capacity, when used for

SRG/SRP applications;

• Cementitious grout well behaved in bonding the steel cords to the masonry substrate and provided an overall

better performance, in terms of ultimate load capacity, than epoxy resin.

ACKNOWLEDGEMENTS

The authors would like to acknowledge Hardwire LLC., Pocomoke City, MD, for providing the steel tapes, the

Department of Architecture and Civil Engineering at the University of Bath for hosting this collaborative

research and the Department of Civil Protection “Consorzio ReLUIS” for supporting this collaborative research.

NOTATION

The following symbols are used in this paper:

M = design bending moment;

MRd,mc = moment capacity for masonry crushing;

MRd,rr = moment capacity for reinforcement rupture;

N = design axial load;

Qexp = experimental load capacity;

Qtheor = theoretical load capacity;

R = radius of curvature;

V = shear action;

VRd = sliding resistance;

breinf = width of the reinforcement;

fM,u = masonry crushing stress;

freinf,u = reinforcement rupture stress;

l = width of the cross section;

t = height of the cross section;

x = rectangular stress-block depth;

ρ = reinforcement area fraction;

σ = normal stress;

σRd = transversal tensile strength for debonding failure;

τ = tangential stress; and

μ = masonry friction coefficient.

REFERENCES

Borri, A., Castori, G. (2004). “Influence of bonding defects in masonry vaults and arches strengthened at their

intrados with FRP.” Proc., Mechanics of masonry structures strengthened with FRP, L. Ceriolo and V. Zerbo,

eds., Venice, Italy, 7-16.

Borri, A., Castori, G., Grazini, A. (2007). “Seismic upgrading of historical masonry buildings with steel

reinforced grout (SRG).” Proc., Fiber-Reinforced Polymer Reinforcement for Concrete Structures (FRPRCS-8),

T. C. Triantafillou, ed., Patras, Greece.

Boughton, B.W., Falconer, R.E., Sumon, S. (1997). “Masonry arch repairs and strengthening”, Proc., Bridge

Modifications 2, stronger and safer bridges, B. Pritchard, ed., Institution of Civil Engineers (Great Britain).

London, UK, 122-130.

Casadei, P., Nanni, A., Alkhrdaji, T., Thomas, J. (2005). “Behavior of Double-T Prestressed Beams

Strengthened With Steel Reinforced Polymer.” Advances in Structural Engineering (ASE), 8(4), 427-442.

Cancelli, A. N., Aiello, M. A., Casadei, P. (2007). “Experimental investigation on bond properties of SRP/SRG

– Masonry systems.” Proc., Fiber-Reinforced Polymer Reinforcement for Concrete Structures (FRPRCS-8), T.

C. Triantafillou, ed., Patras, Greece.

Day, W. (1995). “Strengthening and repair of listed masonry arch bridges.” Proc. First International Conf.,

Arch Bridges, C. Melbourne, ed., Bolton, UK, 567-578.

Fédération International du Béton (FIB) (2001). Externally bonded FRP reinforcement for RC structures,

Technical Bulletin No. 14, CEB-FIB, Lausanne, Switzerland.

Foraboschi, P. (2001). “Strength assessment of masonry arch retrofitted using composite reinforcements.” J. B.

Masonry Soc., 15(1), 17-25.

Foraboschi, P. (2004). “Strengthening of masonry arches with fiber-reinforced polymer strips.” J. Compos.

Constr., 8(3), 96-104.

Hardwire LLC. (2002). “What is Hardwire.” Technical Sheet, <http://www.hardwirellc.com> (Mar. 2, 2008).

Heyman, J. (1982). The masonry arch, Ellis Horwood-Wiley, West Sussex, UK.

Huang, X., Birman, V., Nanni, A., Tunis, G. (2005). “Properties and potential for application of steel reinforced

polymer and steel reinforced grout composites.” Composites, Part B, 36(1), 73-82.

Matana, M., Galecki, G., Maerz, N., Nanni, A. (2005). “Concrete substrate preparation and characterization

prior to adhesion of externally bonded reinforcement.” Proc., Bond Behaviour of FRP in Structures (BBFS

2005), International Institute for FRC in Construction, Hong Kong, China, 133-139.

Prota, A., Tan, K. Y., Nanni, A., Pecce, M., Manfredi, G. (2006). “Performance of RC shallow beams externally

bonded with steel reinforced polymer.” ACI Structural Journal, 108(2), 163-170.

Toray Industries (2003). “T700S.” Technical Sheet, <http://www.soficar-carbon.com>, (Nov. 21, 2007).

Triantafillou, T.C. (1998). “Strengthening of masonry structures using epoxy-bonded FRP laminates.” J.

Compos. Constr., 2(2), 96-104.

Valluzzi, M.R., Valdemarca, M., Modena, C. (2001) “Behaviour of brick masonry vaults strengthened by FRP

laminates.” J. Compos. Constr., 5(3), 163-169.

Wobbe, E., Silva, P., Barton, B. L., Dharani, L. R., Birman, V., Nanni, A., Alkhrdaji, T.,Thomas, J., Tunis, G.

(2004). “Flexural capacity of RC beams externally bonded with SRP and SRG.” Proc., Soc. Advancement of

Materials and Process Engineering (SAMPE 2004), Long beach, CA.

Table 1: Mechanical properties of the reinforcement. Reinforcement Type Property

Steel cords (Type 1) Steel cords (Type 2) Carbon fibers Cross sectional area (mm2/mm) 0.098 0.383 0.167

Tensile load (N/mm) 242 635 817 Elastic Modulus (N/mm2) 160000 143000 230000

Ultimate strain (%) 1.6 1.2 2.1

Table 2: Mechanical properties of the matrices. Matrix Type Property

Polymeric resin Cementitious grout Tensile strength (N/mm2) 32 3.6 Flexural strength (N/mm2) > 35 6.6 Elastic Modulus (N/mm2) 10000 -

Bonding, ASTM D 4541 (N/mm2) > 3.5 > 2.5

Table 3: Test Matrix. Composite type

Specimen Position of reinforcement Matrix Reinforcement

No. of plies Boundary conditions

UN.01 - - - - - EX.01 Cementitious grout Steel cords (Type 2) 1 - EX.02 Polymeric resin Steel cords (Type 2) 1 - EX.03 Cementitious grout Steel cords (Type 2) 1 Steel anchors: angle plates EX.04

Extrados

Polymeric resin Carbon fibers 1 - IN.01 Cementitious grout Steel cords (Type 1) 1 - IN.02 Polymeric resin Steel cords (Type 1) 1 - IN.03 Cementitious grout Steel cords (Type 1) 2 Steel anchors: flat plates IN.04

Intrados

Polymeric resin Carbon fibers 1 - Extrados Steel cords (Type 2) 1

IN+EX.01 Intrados

Cementitious grout Steel cords (Type 1) 1

-

Table 4: Comparison among the experimental results.

Specimen Ultimate load

capacity (kN)

Load point deflection

(mm)

Reinforcement strain (με)

Springing thrust (kN)

Mode of failure

UN.01 0.7 1.5 - 0.2 Mechanism EX.01 9.2 32.53 3598 8.5 - EX.02 13.6 52.39 2759 13.4 Shear sliding EX.03 23.4 49.60 8166 23.8 Shear sliding EX.04 11.5 44.57 6236 10.7 Shear sliding IN.01 16.2 28.60 10723 5.1 Laminate rupture IN.02 14.7 28.74 4219* 4.7 Laminate rupture IN.03 21.7 34.46 10533 6.3 Laminate debonding IN.04 12.3 11.40 2712* - Laminate debonding

IN+EX.01 33.0 34.82 6503 3.9 Shear sliding + Laminate debonding

* Gauge produced unreliable data from this point

Table 5: Theoretical results.

Specimen Analytical model:

load capacity (kN)

Analytical model:error of the model

(Qtheor/Qexp)

Analytical model: mode of failure

UN.01 0.73 1.03 Mechanism EX.01 12.1 1.32 Shear sliding EX.02 12.1 0.89 Shear sliding EX.03 22.7 0.97 Shear sliding EX.04 12.1 1.05 Shear sliding IN.01 15.1 0.93 Laminate rupture IN.02 12.4 0.84 Laminate rupture IN.03 23.8 1.09 Laminate rupture IN.04 30.4 2.47 Laminate rupture

IN+EX.01 31.0 0.94 Shear sliding

a) b) Figure 1: Thrust line and static scheme of arches strengthened at: a) extrados; b) intrados.

a) b) Figure 2: Adhesion tests: a) direct shear tests; b) pull-off tests.

a) b) Figure 3: a) Direct shear tests: SRP specimens failure mode; b) Pull-off tests: SRG specimens failure mode.

Figure 4: Test Setup.

a) b) Figure 5: Application of steel anchors: a) angled plates (80×120×200 mm); b) flat plates (50×200×1.2 mm).

a) b) Figure 6: Shear sliding along a mortar joint: a) first joint closest to the springer (arch EX.02); b) first joint closest to the edge of the steel anchor (arch EX.03).

a) b) Figure 7: a) Laminate rupture: arch IN.02; b) Laminate debonding: arch IN.03.

Figure 8: Shear sliding and laminate debonding: arch IN+EX.01.

Figure 9: Load-deflection curves (measured at the location where load was applied).

Figure 10: Load-strain curves.

Figure 11: Load-springing thrust curves.