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Editor’s quick points

n This paper describes the structural behavior of precast,prestressed concrete sandwich wall panels reinforced with a

carbon-fiber-reinforced polymer (CFRP) shear grid to achieve

composite action.

n Use of CFRP as a shear transfer mechanism was intended to

increase thermal insulation efficiency, enhance service life, and

increase overall structural capacity

n Test results of the experimental program were compared with

theoretical predictions of fully composite and noncomposite

actions to evaluate the percent composite action and to assess

the optimum panel configuration.

Behavior

of precast,

prestressed

concrete

sandwich

wall panels

reinforced

with CFRP

shear gridBernard A. Frankl,Gregory W. Lucier, Tarek K. Hassan,and Sami H. Rizkalla

Precast, prestressed concrete sandwich wall panels are

typically used for building envelopes. Such panels con-sist of two outer layers of precast, prestressed concrete

separated by an inner layer of insulation. The panels can

support gravity loads from floors or roofs, resist normal

or transverse lateral wind loads, insulate a structure, and

provide interior and exterior finished wall surfaces. Typical

panels are fabricated with heights up to 45 ft (14 m) and

with widths up to 12 ft (3.7 m). Concrete wythe thick-

nesses range from 2 in. to 6 in. (50 mm to 150 mm) with

overall panel thicknesses ranging from 5 in. to 12 in. (130

mm to 300 mm).

Precast, prestressed concrete sandwich wall panels maybe designed as noncomposite, partially composite, or fully

composite. Defining and designing for a partial degree of

composite action can significantly increase the structural

efficiency and reduce both initial and life-cycle costs of

a panel, compared with the fully noncomposite case. The

degree of composite action depends on the nature of the

connections between the two concrete wythes. Commonly

used shear transfer connectors include wire trusses, bent

wires, and solid zones of concrete penetrating the insula-

tion wythe (Fig. 1). Increasing the degree of composite ac-

tion between wythes increases the structural capacity of a

given panel, making it more structurally efficient. However,

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Lee and Pessiki found that a three-wythe panel with stag-

gered longitudinal solid concrete zones exhibits behavior

similar to that of a fully composite panel.6 They also

observed that transfer of the prestressing force induced

cracks in the concrete wythes parallel to the prestress-

ing strands. A finite-element analysis was conducted to

investigate the prestressing forces during release. Results

of the analysis showed that modeling the concrete and the

insulation with solid block elements provided results close

to the measured values.

Salmon et al.7 introduced the use of fiber-reinforcedpolymer (FRP) bars formed in a truss orientation in place

of metal wire trusses. Test results showed that the use

of FRP achieved a high level of composite action and

provided thermal benefits similar to noncomposite precast,

prestressed concrete sandwich wall panels. Following the

same concept, a carbon-fiber-reinforced polymer (CFRP)

shear connection grid was used in the construction of pre-

cast, prestressed concrete sandwich wall panels in 2003.8

Because carbon fibers have a thermal conductivity that is

about 14% that of steel, connecting concrete wythes with

carbon grid allows a panel to develop composite structuralaction without thermal bridges. Therefore, the insulating

value of the panel is maintained.1 The grid was oriented di-

agonally between the concrete wythes, normal to the wall

surface, allowing a truss mechanism to develop.

This paper describes the behavior of six full-scale precast,

prestressed concrete sandwich wall panels. The panels

were composed of two outer precast, prestressed concrete

wythes and an internal layer of insulation with shear grid

reinforcement placed through the core into each concrete

wythe. The various parameters considered in the current

study included the type of insulation, presence of solid

traditional composite shear connections have the nega-

tive consequence of thermally bridging the two concrete

wythes, thus decreasing the thermal efficiency.

Wall panels were first introduced during the 1960s as

double-tee sandwich panels.1 Solid concrete zones were

used between the double-tee and the inner concrete wythe

to develop composite action. Double-tee sandwich panels

provided a robust structural wall but sacrificed the poten-

tial thermal savings. Flat concrete slabs were soon used in

place of double-tees to reduce the thickness of the building

envelope and to improve the aesthetics of a structure. Asin double-tee sandwich panels, composite action between

the wythes of flat slab sandwich wall panels was often

achieved through solid concrete zones.

More recently, steel ties and wire trusses were introduced

to replace solid concrete zones. Steel wythe connections

improved the thermal performance of sandwich wall panels

compared with solid concrete zones, but such ties still act

as thermal bridges.1 Noncomposite panels were introduced

in the 1980s and aimed at addressing the thermal deficien-

cies created by steel ties. Noncomposite panels contain

minimal shear connectors to substantially reduce thepotential for thermal bridging but sacrifice the structural

efficiency of a composite structure.

Despite their lower structural capacity, noncomposite

panels became popular due to their thermal savings and

architectural characteristics. The typical design method for

precast, prestressed concrete sandwich wall panels often

assumes noncomposite action.2 In practice, however, panels

generally exhibit partially composite behavior. Test results

by several researchers have shown that significant shear

transfer occurs between the wythes.3–5

Figure 1. Commonly used shear transfer connectors include wire trusses, bent wires, and solid zones of concrete penetrating the insulation core.

Note: CFRP = carbon-fiber-reinforced polymer.

Wire truss connector Bent wire connectors Solid concrete zone CFRP grid material sample CFRP grid shear transfer mechanism in section cut

from a tested pane

(insulation removed)

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(Fig. 2). All panels were 8 in. (200 mm) thick and consist-

ed of three layers. Table 1 summarizes the configurationsof the tested panels.

Panels EPS1, EPS2, XPS1, XPS3, and XPS4 consisted of

two 2-in.-thick (50 mm) concrete wythes with a 4-in.-thick

(100 mm) insulation layer in between. This arrangement

was designated as a 2-4-2 panel configuration. One wythe

of a 2-4-2 panel included two 2-in.-thick (50 mm) and

24-in.-wide (610 mm) internal pilasters along the full

height of each panel at 1 / 4 and3 / 4 widths (Fig. 3). The two

pilasters were provided to carry axial loads from the two

corbels located at the top of the inner panel face. Lifting

anchors on the inner face were centered on the pilasters in2-4-2 panels, so these anchors did not bridge the concrete

wythes. Two lifting anchors were also included on the

top edge of each panel. These anchors spanned between

concrete wythes. Panel XPS2 consisted of a 4-in.-thick

(100 mm) concrete wythe, a 2-in. (50 mm) layer of insula-

tion, and an outer 2-in. (50 mm) concrete wythe. Figure 4

shows this configuration, which was designated 4-2-2 with

two corbels located at the top of the 4-in.-thick wythe. The

4-2-2 panel was designed to carry the axial load through its

thicker concrete wythe and therefore did not have internal

pilasters.

concrete zones, panel configuration, and shear grid rein-

forcement ratio.

The loading sequence for each panel was selected to simu-

late the effect of service gravity and wind loads for a 50-

year lifespan. Load and support conditions were designed

to mimic field conditions. Test results from the experimen-

tal program were compared with theoretical predictions

to evaluate the percent composite action achieved by each

tested panel.

Experimental program

Six precast, prestressed concrete sandwich wall panelswere designed and tested to evaluate their flexural response

under combined vertical and lateral loads. The study

included panels fabricated with two different insulation

types: expanded polystyrene (EPS) insulation and extruded

polystyrene (XPS) insulation. According to the manufac-

turer, the selected EPS insulation had a nominal density of

1 lb/ft3 (16 kg/m3) and a nominal compressive strength of

13 psi (90 kPa). The selected XPS insulation had a nominal

density of 1.8 lb/ft3 (29 kg/m3) and a nominal compressive

strength of 25 psi (170 kPa).

The panels were 20 ft tall × 12 ft wide (6.1 m × 3.7 m)

Inner panel view during testing Outer panel view during testing

Figure 2. The panels were 20 ft tall × 12 ft wide (6.1 m × 3.7 m).

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CFRP shear grid was provided between the two concrete

wythes to transfer the shear stresses across the insulation

and to develop a composite action between the wythes.

The CFRP grid was placed in strips running parallel to the

longitudinal axis of each panel (Fig. 5). The CFRP strips

were embedded 3 / 4 in. (19 mm) into each concrete wythe.

All panels except XPS4 contained the same grid layout.

Panel XPS4 contained an additional 30 ft (9.1 m) of shear

grid. In addition to the CFRP grid, panel XPS1 contained

Each concrete wythe was reinforced with a sheet of

welded-wire reinforcement in the plane of the wythe and

prestressed in the longitudinal direction by five 270 ksi

(1860 MPa), low-relaxation prestressing strands. The

diameters of the prestressing strands in the 2-in.-thick (50

mm) and 4-in.-thick (100 mm) concrete wythes were 3 / 8 in.

(10 mm) and 1 / 2 in. (13 mm), respectively. Figures 3 and 4

show the strands used and their initial tension levels.

Figure 3. The configuration and dimensions of 2-4-2 panels. Panel in photo cut to show cross section. Note: CFRP = carbon-fiber-reinforced polymer; WWR = welded-wirereinforcement.

Table 1. Summary of experimental tests and results

Panel

identificationInsulation

Wythe

thicknesses, in.

Solid

zones

Shear grid

layout

Concrete

strength, psi

Failure load

(1.2D + 0.5L r +…)

Service load

deflection

D + L r + W

EPS1 EPS 2-4-2 No Layout 1 7620 2.8W 120 h /460

EPS2 EPS 2-4-2 No Layout 1 7670 1.8W 150 (2.8W 120) h /500

XPS1 XPS 2-4-2 Yes Layout 2 10,080 1.6W 120 h /1480

XPS2 XPS 4-2-2 No Layout 1 8790 3.2W 120 h /755

XPS3 XPS 2-4-2 No Layout 1 7670 0.7W 120 n.a.

XPS4 XPS 2-4-2 No Layout 3 7340 1.8W 120 h /700

Note: D = dead load; h = height = 240 in.; L r = roof live load; n.a. = not applicable; W = wind load at a selected design wind speed; W 120 = 6.96 kip;

W 150 = 10.56 kip. 1 in. = 25.4 mm; 1 kip = 4.448kN; 1 psi = 6.895 kPa.

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10 discretely located solid concrete zones throughout the

height and width of the panel (Fig. 5).

Test setup

All panels were tested in the laboratory using a steel testing

frame that allowed for simultaneous application of gravity-

and lateral loads (Fig. 2). Reverse-cyclic lateral loads were

applied at a rate of 1 cycle per 10 sec (0.1 Hz) to simulate

wind pressures. The testing frame consisted of one braced

frame on each side of the panel to support an upper cross

beam. This cross beam in turn provided the upper lateralsupport to the panel. The entire setup was anchored to the

laboratory strong floor. A closed-loop hydraulic actuator,

supported by a strong reaction wall, applied the lateral

load.

Each panel was simply supported in the testing frames at

the top and bottom edges. The bottom of the panel was

supported by a hinge, which restrained horizontal and

vertical movements while allowing for rotation. The center

of the hinge was located 1 in. (25 mm) below the bottom

of each panel. The top of each panel was supported using a

sliding pin connection that restrained horizontal motion but

allowed for vertical movement and rotation. The center of

the sliding pin was located 4 in. (100 mm) above the top of

the panel.

Vertical loads were applied to the top of each corbel by a

hydraulic jack and cable (Fig. 2) to simulate the effects of

a 60-ft-span (18.3 m) double-tee roof system. The jack was

connected to an accumulator to maintain a constant axial

load as the panel deformed. Lateral loads were applied by

the actuator connected to a spreader beam system, which

was used to push and pull the panel to simulate wind pres-

sure and suction. Two loading tubes were provided at each¼-height of each panel, one on each wythe, to distribute

the lateral load across the width of the panel. The lateral-

loading mechanism included a vertical spreader beam

that could shorten and elongate as the panel deformed to

prevent the transfer of any unintended forces to the panel.

Each panel was subjected to reverse-cyclic loading begin-

ning at a level equivalent to 70% of the service load. The

loading regime was selected using a Weibull distribution to

simulate wind loads over a 50-year service life.9

All panels were instrumented to measure lateral deflection,

relative displacement between the two concrete wythes,

Figure 4. The configuration in this drawing was designated as 4-2-2 with two corbels located at the top of the 4-in.-thick wythe. Panel cut to show cross section. Note:

CFRP = carbon-fiber-reinforced polymer; WWR = welded-wire reinforcement. 1 in. = 25.4 mm; 1 ft = 0.305 m; 1 lb = 4.448 N.

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surface strain of the concrete, and applied axial and lateral

loads. The strain profile across the thickness of each panel

was measured with four electrical-resistance strain gauges

across the panel section at three locations along the height.

Each panel was subjected to 3710 fully reversed lateral-

load cycles at 45% of the factored lateral wind load, equiv-alent to 0.7W , with a factored axial load of 1.2 D + 0.5 Lr in

place, where W is wind load, D is dead load, and Lr is roof

live load. The initial cycles were followed by 177 cycles at

50% of the factored lateral wind load (0.8W ) with the fac-

tored axial load applied. Subsequent individual cycles were

applied at 60%, 80%, and 100% of the factored lateral

wind load (1.0W , 1.3W , 1.6W ), all with the factored axial

load applied. After completion of all lateral cycle loads in

the presence of gravity load, the lateral load was increased

in one direction only until failure.

Results and discussion

Lateral displacements

Figure 6 depicts the measured lateral displacement at

midheight for the different panels. In general, the measured

lateral deflections due to the applied axial load only were

found to depend on the thickness of the panel configuration

(2-4-2 or 4-2-2) and also on the type and configuration of

shear transfer mechanism used. Lateral deflection due to

axial load alone was shown as the offset deflection at zero

lateral-load level.

The allowable displacement of h /360 (where h is the height

of the panel) at service load level (as per the American

Concrete Institute’s ACI 533R-93, Guide for Precast

Concrete Wall Panels10) was compared with the measured

values for the tested panels. The two EPS-insulation panels

EPS1 and EPS2 behaved almost identically throughout the

loading cycles. The panels’ stiffnesses remained constantto a lateral load of 15 kip (67 kN), or 2.2W 120 (where W 120

is the wind load at a wind speed of 120 mph), beyond

which concrete cracking occurred, considerably reducing

the stiffness (Fig. 6). Similar behavior was observed for

XPS2. The behavior of panels XPS1 and XPS4 remained

linear to a lateral load level of about 10 kip (44.5kN), or

1.4W 120. Panel XPS3 failed prematurely at a lateral load

level of 5 kip (22 kN).

The maximum measured displacements at service-load lev-

el for EPS1 and EPS2 were equivalent to h /460 and h /500,

respectively. These service-level displacements were wellwithin the ACI 533R limit. Among the XPS-insulation

panels, XPS1 experienced the least stiffness degradation

with increased load cycles because of the presence of solid

concrete zones connecting the inner and outer wythes. The

maximum lateral displacement at service-load level for

XPS1 was equivalent to h /1480.

Panels XPS2 and XPS4 exhibited minimal lateral-load

degradation. The measured lateral displacements for both

panels did not increase noticeably throughout the fatigue

cycles. The maximum lateral displacements at service-

load level for XPS2 and XPS4 were equivalent to h /755

Figure 5. Layout of CFRP shear grids. Note: CFRP = carbon-fiber-reinforced polymer. 1 in. = 25.4 mm; 1 ft = 0.305 m.

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Figure 6. Measured lateral displacement at mid-height for the different panels.

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XPS4 showed a clear reduction in composite action with

increasing lateral load. A significant discontinuity in the

strain at ultimate load was observed for these panels, indi-

cating a partial composite behavior at ultimate.

Failure modes

The observed failure modes for EPS1, XPS1, and XPS2

were localized at the tops of the panels in the corbel zones.

Failure was characterized by a shear failure around the cor-bels extending down about 2 ft (600 mm). Figure 8 shows

an example of these failures, which were accompanied by

separation of the top of the panel. Panel EPS2 exhibited a

flexural-shear failure across the width of the panel at about7 / 8 the panel’s height. Panels XPS3 and XPS4 exhibited a

flexural-shear failure across the width of the panel at about7 / 8 the panel’s height along with a simultaneous top-of-

panel separation.

All panels with sufficient shear transfer mechanisms

exhibited deflections well below the limiting value speci-

fied by ACI 533R and sustained loads prior to failure inexcess of their factored design loads (Table 1). However,

panel XPS3, which had a gap between the outer wythe and

foam, failed prematurely prior to the service load with high

deflections.

Uniform design pressures for panels EPS1, XPS1, XPS2,

XPS3, and XPS4 were assumed to be 29 lb/ft2 (1.4 kPa),

corresponding to a design wind speed of 120 mph (193

km/hr). Although designed for 29 lb/ft2 (1.4 kPa), EPS2

was tested for an equivalent pressure of 44 lb/ft2 (2.1 kPa),

corresponding to a design wind speed of 150 mph (241

km/hr). Thus, the lateral fatigue loading on EPS2 was

and h /700, respectively. Failure of panel XPS3 occurred

before reaching the design service-load level. Test results

suggested that the accumulated degradation for XPS3 was

substantial compared with other panels. A gap between

the outer concrete wythe and the foam, observed before

testing, likely contributed to the premature failure. This

gap measured about 1 / 4 in. (6 mm) and was visible along

the majority of both 20 ft (6.1 m) panel sides. The gap was

identified by the precast concrete producer after stripping

the forms, but because the panel was intended for testing, itwas decided not to reject the piece.

Strain profiles

The strain profiles across the thickness of the panels were

measured to determine the degree of composite action

between the two concrete wythes. Figure 7 shows typical

results recorded during the testing of the panels at ultimate

load. For all panels, the inner wythe experienced com-

pressive strains while the outer wythe experienced tensile

strains under the effect of applied factored gravity loads.

Test results showed that EPS-insulation panels EPS1 andEPS2 as well as XPS1 with solid concrete zones exhib-

ited and maintained a high level of composite action until

failure.

The strain profile at ultimate load indicates that the neutral

axes of these panels were located closer to the elastic cen-

troid of the composite cross section rather than the elastic

centroid of each individual wythe. The measured strains for

XPS2 (4-2-2 configuration) indicated that each wythe acted

independently in carrying the applied loads and the neutral

axis was located within the thickness of each wythe. This

behavior indicated noncomposite action. Panels XPS3 and

Figure 7. Strain profile distribution for different panels at ultimate loads. Note: The negative (-) sign indicates compression. Note: 1 kip = 4.448 kN.

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M cr = cracking moment

M a = acting moment

I g = gross moment of inertia of the section

I cr = cracked transformed moment of inertia of the section

Recent research findings by Bischoff and Scanlon have

demonstrated that Eq. (2) is only applicable for flexural

members with an I g / I cr ratio less than 3, which corresponds

to beams or slabs with a steel reinforcing ratio greater than

1%.12 It has been demonstrated that applying Eq. (2) tocross sections with an I g / I cr ratio greater than 3 consider-

ably overestimates the member stiffness.11 Equation (3)

is an alternative formulation of the effective moment of

inertia that was proposed by Bischoff and Scanlon.11

1

I

M

M

I

I

I I

1

ef f

a

cr

g

cr

cr

g2 #=

- -f p > H (3)

The ratios of I g / I cr for the inner and outer wythes of thetested panels ranged from 20 to 183, which is significantly

higher than the limiting value of 3 proposed by Bischoff

and Scanlon.11 Figure 9 shows comparisons between the

measured and predicted displacements for different panels

using both approaches for the effective moment of inertia.

For the theoretical fully composite case, applied loads and

moments were assumed to act on the fully composite sec-

tion. For the theoretical noncomposite case, it was assumed

that the applied axial load was resisted by the inner wythe

(with corbels) alone. Applied moments were assumed to

be distributed to the inner and outer wythes depending on

higher than the load used for the other panels.

Analysis of sandwichwall panels

To evaluate the degree of composite action for the panels,

the measured lateral displacements were compared with the

predicted values assuming both fully composite and fully

noncomposite behaviors. The percentage of composite ac-

tion k was evaluated for all tested panels using Eq. (1).

k 100exp

noncomposite

n on co mp os it e e rim en ta l

compositeD D

D D=

-- a k (1)

where

∆experimental = measured displacement at a selected load level

∆composite = corresponding theoretical displacement as-

suming fully composite behavior

∆noncomposite = corresponding theoretical displacement as-

suming fully noncomposite behavior

To determine theoretical panel displacements beyond

cracking, the effective moment of inertia I eff was calcu-

lated by Eq. (2) in accordance with ACI’s Building Code

Requirements for Structural Concrete (ACI 318-05) and

Commentary (ACI 318R-05).11

1 I M

M

M

M I I I

ef f

a

cr

a

cr

g cr g

33

#= + -f fp pR

T

SSSS

V

X

WWWW

(2)

where

Figure 8. Typical observed failure modes for EPS1, XPS1, and XPS2.

Corbel-zone shear failure on inner wythe Outer wythe cracking following top-of-panel separation at failure

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for XPS2 was significantly lower than for other panels,the maximum displacement at service load level was still

within the ACI 533R limitations. Panel XPS3 failed prior

to reaching the combined axial and lateral service-load

condition.

Conclusion

The flexural behaviors of six full-scale insulated precast,

prestressed concrete sandwich wall panels were investi-

gated. The panels were subjected to monotonic axial and

reverse-cyclic lateral loading to simulate gravity and wind

pressure loads, respectively. Based on the findings of this

their individual stiffnesses.

The distribution ratio was calculated using the average of

their gross and cracked moments of inertia. The percentage

of composite action was calculated for each panel under

the combined axial and lateral service load, and Table 2

summarizes the results. In all calculations, a 20 ft (6.1 m)

nominal span was assumed.

The estimated composite action for EPS1, EPS2, XPS1,

and XPS4 was nearly 100%. Panel XPS2 exhibited 18%

composite action under the combined axial and lateral

service load. Although the percent of composite action

Figure 9. Load-lateral displacement behavior for the panels.

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study, several conclusions were made:

• Panel stiffness and deflections are significantly af-

fected by the type and configuration of the shear

transfer mechanism. Panel stiffness is also affected by

the type of foam.

• Values of percent composite action near 100% can be

achieved with CFRP grid shear connections or with

solid concrete zones.

• Appropriate use of CFRP shear grid can provide an

effective shear transfer mechanism in precast, pre-

stressed concrete sandwich wall panels, as evidencedby the behavior of panels EPS1, EPS2, XPS2, and

XPS4. All panels sustained loads in excess of their

factored design loads and exhibited large deformations

before failure. CFRP grid can provide the required

composite action between wythes using either EPS or

XPS foam.

• Appropriate selection of the CFRP shear-grid quantity

and configuration is critical to achieve optimal struc-

tural performance of a panel. Proper quality control in

production is especially important for composite wall

panels.

• For a given shear transfer mechanism, a higher percent

composite action can be achieved using EPS insula-

tion rather than XPS insulation. Use of XPS insulation

requires an increase of the shear reinforcement ratio

compared with EPS insulation.

Acknowledgments

The authors are grateful for the support of AltusGroup and

the assistance provided by Harry Gleich of Metromont

Corp. and Steve Brock of Gate Precast.

Table 2. Percentage of composite action for different panels at service load level

Panel

identification

Experimental

displacement, in.

Composite

displacement, in.

Noncomposite displacement, in. k , % k , %

Bischoff formula

for I eff

ACI 318 formula

for I eff

Bischoff formula

for I eff

ACI 318 formula

for I eff

EPS1 0.21 0.200 39.6 6.70 100.0 99.8

EPS2 0.24 0.200 39.6 6.70 99.9 99.4

XPS1 0.14 0.087 28.8 2.20 99.8 97.6

XPS2 0.46 0.089 0.5 0.50 17.7 17.7

XPS3 n.a.* 0.087 n.a. n.a. n.a. n.a.

XPS4 0.31 0.087 20.1 3.40 98.9 93.5

*Specimen XPS3 failed at a lateral load level less than the design service load.

Note: Composite displacements differ between EPS and XPS 2-4-2 panels due to different values for concrete elastic modulus. I eff = effective moment

of inertia; k = percentage of composite action; n.a. = not applicable. 1 in. = 25.4 mm.

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Members Containing Steel Reinforcement and Fiber-

Reinforced Polymer Reinforcement. ACI Structural

Journal, V. 104, No. 1 (January): pp. 68–75.

8/18/2019 Frankl Et Al 2010

13/13

C—C—

H

H

–H

–H

Notation

D = dead load

h = height of the panel

I cr = cracked transformed moment of inertia of the

section

I eff = effective moment of inertia

I g = gross moment of inertia of the section

k = percentage of composite action

Lr = roof live load

M a = acting moment

M cr = cracking moment

W = wind load

W 120 = wind load at a wind speed of 120 mph

W 150 = wind load at a wind speed of 150 mph

∆composite = corresponding theoretical displacement as-

suming fully composite behavior

∆experimental = measured displacement at a selected load level

∆noncomposite = corresponding theoretical displacement as-

suming fully noncomposite behavior

About the authors

Bernard A. Frankl graduated with

his MS in civil engineering fromNorth Carolina State University in

Raleigh, N.C., and now works for

the South Dakota Department of

Transportation.

Gregory W. Lucier is the manager

of the Constructed Facilities

Laboratory at North Carolina

State University.

Tarek K. Hassan, PhD, is an

associate professor for the

Structural Engineering Depart-

ment at the Faculty of Engineer-

ing at Ain Shams University, and

a senior structural engineer at Dar

Al Handasa, Cairo, Egypt.

Sami H. Rizkalla, PhD, P.Eng., is

a Distinguished Professor of Civil,

Construction and Environmental

Engineering, director of the

Constructed Facilities Laboratory,and director of the National

Science Foundation Industry/

University Cooperative Research Center at North

Carolina State University.

Synopsis

This paper describes the structural behavior of precast,

prestressed concrete sandwich wall panels reinforced

with carbon-fiber-reinforced polymer (CFRP) sheargrid to achieve composite action. Use of CFRP as a

shear transfer mechanism was intended to increase

the thermal insulation efficiency, enhance the service

life, and increase the overall structural capacities of the

panels.

This study included testing of six full-scale sandwichwall panels, each measuring 20 ft × 12 ft (6.1 m × 3.7

m). The panels consisted of two outer prestressed con-

crete wythes and an inner insulation wythe. The study

included two types of insulation and several shear

transfer mechanisms with different CFRP reinforce-

ment ratios to examine the degree of composite action

developed between the two concrete wythes.

All panels were simultaneously subjected to applied

gravity and lateral loads. Reverse-cyclic lateral loads

simulated the effects of wind pressure and suction. All

panels were subjected to approximately 4000 cycles

of lateral loading with the presence of factored gravity

load. Following each fatigue regime, the lateral loads

were increased until failure was achieved. Test results

of the experimental program were compared with theo-

retical predictions of fully composite and noncompos-

ite actions to evaluate the percent composite action and

to assess the optimum panel configuration.

Keywords

Carbon-fiber-reinforced polymer, CFRP, composite,

insulated wall panel, sandwich wall panel, shear grid.

Review policy

This paper was reviewed in accordance with the

Precast/Prestressed Concrete Institute’s peer-review

process.

Reader comments

Please address any reader comments to journal@pci.

org or Precast/Prestressed Concrete Institute, c/o PCI

Journal, 200 W. Adams St., Suite 2100, Chicago, IL

60606. J

Frankl Et Al 2010

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