Blast Performance of Single-Span Precast Concrete Sandwich Wall...

Blast Performance of Single-Span Precast ConcreteSandwich Wall Panels

Clay Naito, M.ASCE1; Mark Beacraft2; John Hoemann, A.M.ASCE3; Jonathan Shull4;Hani Salim, A.M.ASCE5; and Bryan Bewick6

Abstract: A research program was conducted to assess the capability of conventional non-load-bearing insulated precast concrete exteriorwall panels to withstand blast loadings. Typical construction details from the tilt-up and prestressed concrete industries were examined. Thesensitivity of insulation type, reinforcement, foam thickness, and shear tie type on the flexural resistance was assessed. Forty-two single-spanstatic experiments were conducted on 14 different panel designs. From the results of these experiments, resistance functions and deformationlimits for insulated concrete sandwich panels were determined. The resistance functions were used to develop predictive dynamic models forpanels subjected to blast demands. The models were found to be accurate in comparison to measurements from four full-scale blast det-onations. The findings of the research indicate that both prestressed and non-prestressed insulated concrete wall panels meet current rotationallimits defined by the U.S. Army Corps of Engineers for protective structures. Simplified methods for modeling the pressure-deformationcharacteristics of insulated panels can be used to approximate the peak dynamic deformation and reaction loads. DOI: 10.1061/(ASCE)ST.1943-541X.0001020. © 2014 American Society of Civil Engineers.

Author keywords: Sandwich panels; Prestressed concrete; Composite beams; Dynamic tests; Shock and vibratory effects.

Introduction

The use of insulated precast concrete and insulated tilt-up concretesandwich panels for exterior walls is common practice in theUnited States. This form of construction provides a thermally effi-cient and high-mass wall that enhances the energy efficiency andblast resistance of a building, making it a desirable option formilitary and government facilities. Current design recommenda-tions are very restrictive when using these forms of construction,due in large part to the lack of experimental research data on whichto base recommendations. To address this issue, a research programhas been conducted to assess the performance of conventional in-sulated non-load-bearing exterior wall systems under blast loading.The research examines the static performance of wall panelssubjected to uniform loading and uses these results to estimate

the dynamic response of the panels. The models are validatedthrough a series of full-scale blast detonations.

Research Significance

With continued increase in international terrorism incidents, effortsto design military, government, and high-priority civilian buildingsfor blast resistance are becoming more common. To maximizestructural efficiency while maintaining constructability, design cri-teria for blast demands should account for the latest methods ofconstruction. Non-load-bearing, insulated, reinforced concretesandwich wall panels represent such a system and have beenchosen as the focus of this research. These panels can be con-structed either on-site using tilt-up techniques or off-site in aquality-controlled environment. They are erected into place quicklyand provide both thermal efficiency to the structure and blast pro-tection for the building and contents thereof during and after a blast.

Acceptable performance of reinforced concrete panels subjectedto blast demands is dependent on the level of damage developedunder the expected demands. The U.S. Army Corps of Engineers(USACE 2008) prescribes the damage levels which are tied to re-sponse limits. Limits have been developed for both prestressed andnon-prestressed concrete walls, beams, and slabs subjected toflexural response. Insulated sandwich panels are traditionally de-signed for wind and handling loads and have not been thoroughlyexamined for extreme out-of-plane loading such as that generatedduring an explosion. The objective of this research effort was toexamine the performance of insulated sandwich wall panels underthese demands and to determine acceptable response limits.

Response Limits

The response limit of a concrete component in a blast environ-ment is assessed relative to two variables: deflection ductility andend rotation. Deflection ductility, μ, is defined as the maximum

1Associate Professor, Dept. of Civil and Environmental Engineering,Lehigh Univ. ATLSS Center, 117 ATLSS Dr., Bethlehem, PA 18015(corresponding author). E-mail: [email protected]

2Graduate Student Researcher, Dept. of Civil and EnvironmentalEngineering, Lehigh Univ. ATLSS Center, 117 ATLSS Dr., Bethlehem,PA 18015. E-mail: [email protected]

3Research Civil Engineer, U.S. Army Engineer Research and Develop-ment Center, 3909 Halls Ferry Rd., CEERD-GS-V, Bldg. 5001, Vicksburg,MS 39180-6199; formerly, Air Force Research Laboratory SupportContractor, Applied Research Associates, Inc., Tyndall AFB, FL.

4Structural Engineer, Black & Veatch, Federal Service Division,1805 Meadow Moor Dr., Webb City, MO 64870.

5Associate Professor, Dept. of Civil and Environmental Engineering,Univ. of Missouri, E2509 Lafferre Hall, Columbia, MO 65211-2200.

6Project Engineer, Protection Engineering Consultants, 14144Trautwein Rd., Austin, TX 78737; formerly, Air Force Research Labora-tory, Tyndall AFB, Panama City, FL 32403.

Note. This manuscript was submitted on July 1, 2013; approved onDecember 2, 2013; published online on June 12, 2014. Discussion periodopen until November 12, 2014; separate discussions must be submitted forindividual papers. This paper is part of the Journal of Structural Engi-neering, © ASCE, ISSN 0733-9445/04014096(13)/$25.00.

© ASCE 04014096-1 J. Struct. Eng.

J. Struct. Eng. 2014.140.

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http://dx.doi.org/10.1061/(ASCE)ST.1943-541X.0001020




deflection at midspan,Δmax, resulting from a uniformly distributedload divided by the yield deflection at midspan, Δyield, for thecross-section. The maximum end rotation, θmax, for a simply sup-ported or fixed-fixed beam is approximated relative to theΔmax andthe span length, L, in accordance with Eq. (1). End rotations, θ,as opposed to deflection, are used as a parameter for definingresponse limits to account for the greater flexibility available forlonger-span components

θmax ¼ tan−1½Δmax=ðL=2Þ� ð1ÞThe response limits associate physical response with component

damage and building protection levels (USACE 2008). Five dam-age levels, from Superficial to Blowout, are defined. Each damagelevel respectively corresponds to five levels of decreasing buildingprotection from High to Below Standard. For instance, a buildingwith a High-level of protection would require all primary flexuralprestressed components to have only superficial levels of damage(i.e., deflection ductility less than 1.0). The development of theseresponse limits assumes that the panel is capable of achievinga stable plastic mechanism without loss of strength. To assesswhether these limits are applicable, a series of conventional sand-wich wall panels were experimentally evaluated.

Non-Load-Bearing Sandwich Wall Panels

Concrete sandwich wall panels are widely used across the UnitedStates for construction of building systems. The panels consistof an interior section, or wythe, of insulating foam and exteriorwythes of concrete, as illustrated in Fig. 1. The panel cross-sectionis commonly described relative to the thickness of each wythe ininches starting with the interior wythe (i.e., 3-2-3). The interior andexterior layers are connected with shear ties. Varying the numberand type of shear ties allows the interior and exterior wythes to actas a fully composite, partially composite, or noncomposite system(Naito et al. 2012). The systems can be used in both load-bearingapplications and for non-load-bearing cladding systems. Detailson the design and construction of these systems for conven-tional demands including shipping, handling, wind, and thermalrequirements are summarized in the State of the Art of Precast/Prestressed Concrete Sandwich Wall Panels Report (PCI 2011).

Concrete sandwich walls have a number of properties that makethem a good option for government and military facility construc-tion. The foam sandwich provides a high level of insulation, result-ing in an energy-efficient building envelope. The prefabrication ofthe panels allows for rapid erection of the building and short con-struction schedules. The use of concrete provides a higher inertialmass than other cladding options such as wood or steel framedconstruction. This increased inertial mass provides a greater blastresistance for the facility against external detonations.

To assess the capability of sandwich wall panels to withstandblast loads, single-span panels were examined at quasi-static and

dynamic loading rates. The quasi-static evaluation was conductedto assess the strength and deformation capacity of conventionalpanels. The dynamic study followed to examine the response underdynamic loads and to evaluate the accuracy of predictive methodsin computing response. Fourteen 3.66-m (12 ft) long, 0.81- or0.41-m (32- or 16-in.) wide, single-span wall details were exam-ined under quasi-static demands until failure. Two to three staticexperiments were conducted on each configuration to examine re-peatability, for a total of 42 experiments (Naito et al. 2011a). Fourwall details were examined under dynamic loads generated fromhigh explosive detonations. The dynamic specimens measured3.66 m (12 ft) long and 1.5 m (59 in.) wide. The panels wereexamined under four blast demands. The wythe thickness, insula-tion type, amount and type of reinforcement, and shear tie werevaried among panels. The wall panels were detailed to meet thedesign standards used for the Precast/Prestressed Concrete Industry(PCI 2010) and the methods of the Tilt-Up Concrete Association(TCA 2006). Details of the specimens evaluated in the staticprogram are illustrated in Fig. 2.

Both the PCI and the TCA panels were designed as non-load-bearing in accordance with ACI 318 (ACI 2008). To assess the per-formance of existing construction, the panel designs were based onstandard wind and handling loads with no direct consideration ofblast demands. For the TCA panels, the concrete strength wasspecified as 27.6 MPa (4000 psi). For the PCI panels, the concretestrength was specified as 24.1 MPa (3500 psi) at transfer and34.5 MPa (5000 psi) at 28 days. These strengths are typical of sand-wich panel construction. The higher strengths for PCI representthe increase due to plant fabrication. All prestressing strands are9.5-mm (3/8-in.) diameter, 1862-MPa (270-ksi), low-relaxation,seven-wire strands conforming to ASTM A416 (ASTM 2006a).

Shear Ties and Insulation

The panels are fully insulated; no solid concrete zones are usedbetween wythes. To provide integrity between the interior andexterior concrete faces, shear ties are used. This study examineda number of different shear ties (Fig. 3): (A) THERMOMASSComposite Tie, (B) THERMOMASS Non-Composite Tie, (C) Car-bon Cast C-Grid, and (D) Steel C-Clip. All ties are commerciallyavailable and widely used. Steel ties were included because of theirprevalence in the U.S. The major disadvantage of these ties is thethermal conductivity they possess, which was shown to increaseheat transmittance through a sandwich panel by as much as 20%over glass or carbon fiber composite ties (PCI 2011). The shearproperties of these ties were examined and discussed by Naitoet al. (2012). Three types of insulation commonly used for sand-wich panel construction were examined: expanded polystyrene(EPS), extruded expanded polystyrene (XPS), and polyisocyanu-rate (Polyiso). The thermal resistivity, cost, and density increasedfrom EPS to XPS to Polyiso.

Fig. 1. 3-2-3 sandwich panel



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Experimental Program for Static Tests

Table 1 presents the configurations examined for the panels. Eightpanel details were prestressed (PS) with seven-wire strand, and fourpanel details were reinforced with conventional rebar (non-PS).Three specimens were tested for each configuration. For compositepanels, the number of ties was chosen to provide the comparablelevels of shear strength. The stiffness of the system howevervaries considerably based on the tie structure and insulation used.

The average measured concrete strengths for each panel are shownin Table 1. The concrete compressive strength, f 0

c, was measured inaccordance with ASTM C39 (ASTM 2005). Welded-wire rein-forcement (WWR) conformed to ASTM A82 (ASTM 2007a), andall deformed reinforcement conformed to ASTM A615 Grade 60(ASTM 2012) or ASTM A706 (ASTM 2009).

A static load test setup for wall panels was introduced in Naitoet al. (2011a) and is also illustrated in Fig. 4. The static experimentconsisted of 16 equal point loads applied through a load tree. Thecenter of the load tree was incremented at a rate of approximately0.25 mm=s. The boundary conditions consisted of heavy-walledsteel pipe providing simple supports at a 10-ft (3.05-m) span. Theend rotations, slips, and midspan deflection were measured foreach panel.

Experimental Results of Static Tests

Static Resistance Functions

A representative five-point resistance backbone of each experimentwas developed and used to quantify the average response of the

Fig. 2. Specimen sections for static tests

Fig. 3. Shear tie types

Table 1. Static Test Matrix

Panel IDWythe

configuration InsulationReinforcement

(longitudinal/transverse) Shear tie typef 0c interior wythe,

kPa (ksi)f 0c exterior wythe,

kPa (ksi)

Non-PSTS1 6-2-3 XPS #3/WWR B 33.9 (4.9) 38.0 (5.5)TS2 3-2-3 XPS #3/#3 A 34.0 (4.9) 38.1 (5.5)PCS3a 3-2-3 EPS #5/WWR C 58.0 (8.4) 58.0 (8.4)PCS7 3-3-3 XPS #5/#3 A 61.1 (8.9) 61.1 (8.9)

PSPCS1 3-2-3 EPS 3=8∅ strand/WWR D 56.7 (8.2) 56.7 (8.2)PCS2 3-2-3 EPS 3=8∅ strand/WWR C 57.9 (8.4) 57.9 (8.4)PCS3b 3-2-3 EPS 3=8∅ strand and #5/WWR C 59.6 (8.6) 59.6 (8.6)PCS4 3-3-3 XPS 3=8∅ strand/#3 D 60.8 (8.8) 60.8 (8.8)PCS5 3-3-3 XPS 3=8∅ strand/#3 A 60.6 (8.8) 60.6 (8.8)PCS6 3-3-3 XPS 3=8∅ strand/WWR C 59.8 (8.7) 59.8 (8.7)PCS8 3-3-3 Polyiso 3=8∅ strand/#3 A 60.8 (8.8) 60.8 (8.8)PCS9 3-3-3 Polyiso 3=8∅ strand/WWR C 59.4 (8.6) 59.4 (8.6)



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specimens. A procedure was developed to minimize the errorbetween the energy of the measured pressure versus rotationperformance and that of the resistance function. The five pointscorresponded to the maximum pressure, Pmax, point M, secantstiffness to point K, and two post-peak levels, S and T, shownin Fig. 5. Point T represented blowout and was defined as the pointat which 50% of the maximum pressure resistance remained. Thepoints K and S were chosen as a percentage of Pmax that wouldminimize the error between the measured pressure-rotation re-sponse and the final resistance functions. The error was computedas shown in Eq. (2):

error ¼RPθdθðactualÞ − R

PθdθðbackboneÞRPθdθðactualÞ ð2Þ

The value of K and S were determined as 64% Pmax and72% Pmax, respectively. These values provided the lowest averageerror on the entire data set of one-span panel tests. Table 2 showsthe average values for the four non-zero points of the resistancefunction for each panel configuration. Figs. 6 and 7 graphicallyshow the non-prestressed and prestressed resistance functions,respectively.

Static Results Summary

Table 3 summarizes the current design limits (USACE 2008) andthe average actual responses for non-prestressed and prestressedconcrete insulated elements subject to flexure without tensionmembrane action. Response limits are defined in terms of thecomponent damage achieved. Damage increases from no visiblepermanent damage, some repairable permanent deflection, signifi-cant unrepairable permanent deflection, to component has failed.These damage levels are designated as Superficial, Moderate,Heavy Damage, and Hazardous Failure, respectively. The mea-sured response for Superficial, Moderate, Heavy, and Hazardouslevels of damage correspond with points L, M, S, and Tof the mea-sured resistance functions. The end of the Moderate damage regioncorresponds to the point of maximum resistance. This is in line withthe designation of some repairable damage in the component.

The average measured performance for both non-prestressedand prestressed panels exceeded the acceptable response limitsin all but one case. The exception was for non-PS panels thatdid not on average meet the 10.0° Hazardous damage level limit.This was attributed to the loss of composite action due to failure ofthe ties at large deformations, which limited their ability to maintainlarge rotations. Further details on loss of composite action aredetailed in Naito et al. (2012).

The results of the static tests are presented in Table 4. Mean andstandard deviation values for each specimen configuration are re-ported. The strengths are compared with estimated composite andnoncomposite flexural capacities calculated with as-built dimen-sions. The flexural strength was estimated, accounting for steel inonly the tension wythe. Composite strength was computed assum-ing a linear strain profile across the entire cross-section depth. Thenoncomposite strength was computed assuming each wythe actedindependently (Fig. 8). Due to the presence of shear ties in thepanels, strength varied from noncomposite to fully composite. Thepercentage of fully composite was computed from the measuredstrengths, using Eq. (3):

ðMeasuredmoment capacityÞ − ðEstimated noncomposite capacityÞðEstimated composite capacityÞ − ðEstimated noncomposite capacityÞ ¼ %Composite ð3Þ

The panels range from a low of 49% composite to 185%composite with an average of 95% composite. In general, theprestressed panels were not able to reach their estimated compositeflexural strength, while the majority of reinforced concrete panelsdid achieve their composite capacity.

The ultimate strength of the panels was limited by a localizedflexural mechanism occurring near midspan. At elevated demands,the ties deform and fracture, and due to the compressibility of thefoam, the panels tended to fold. The strength associated with thisbehavior can be bounded by the assumption that the foam is com-pletely crushed and the interior and exterior wythes are touching.A folding moment capacity was estimated based on this experimen-tal observation, similar to the composite estimation, by assigninga zero thickness to the foam wythe. The concept is illustrated inFig. 8 and summarized in Table 4. The folding capacity provideda conservative estimate of capacity for all cases except for the over-reinforced panel PCS3b.

Experimental Observations of Static Performance

The responses of the panels were sensitive to the shear tie connec-tors used (Fig. 9). This is apparent when comparing PCS1 andPCS2, which contain the same flexural reinforcement. PCS1 usedsteel flexible C-clips, while PSC2 used a stiff C-grid. The C-gridpanel exhibited a higher stiffness and strength but achieved lessductility than the C-clip reinforced panel. This is, again, illustratedin a comparison of PCS4, 5, and 6, which contained the same flexu-ral reinforcement and XPS insulation. The ties varied and includedC-clips, GFRP pins, and C-grid for PCS4, 5, and 6, respectively.The postcracking stiffness varied in accordance with the tie stiff-ness. The C-clip was the lowest, followed by the glass pins andthe C-grid.

The responses of the panels were sensitive to the foam typeused. A 3-2-3 and a 3-3-3 wythe configuration panel were testedwith the same reinforcement. Based on composite action, the 3-3-3panel should have provided a higher moment capacity than that of

Fig. 4. Loading tree schematic



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the 3-2-3 panel, due to the larger moment arm. The two panels,however, achieved approximately the same strength, as revealedby comparing PCS1 to PCS4 and PCS2 to PCS6. In both cases,the use of XPS (over EPS) insulation resulted in a softer postcrack-ing stiffness, larger ductility, and lower strength (even with thelarger moment arm). The use of unsealed Polyiso increased thestrength of the panels further, as seen by comparing PCS6 andPCS9. It was noted that both the EPS and Polyiso had a roughabsorbent surface compared to the XPS insulation, which wassmooth and sealed. The additional roughness provided additionalhorizontal shear strength against flexural demands. The use of oneinsulating foam layer was critical for composite action. PCS8 wasfabricated using two 3.8-cm layers of Polyiso with a sealed inter-face between layers. This break in the insulation resulted in areduction in strength and stiffness from that of XPS, as seen bycomparing PCS5 and PCS8 (Fig. 7).

The use of PS strand versus conventional non-PS reinforcementis illustrated in a comparison of PCS2 and PCS3a. Both specimensachieved approximately the same flexural strength. The use of non-PS reinforcement resulted in a softening behavior after cracking.This was not as evident in the PS section (PCS2). The use ofPS reduced the deformation ductility by approximately 50%. Thisinability to support larger deformations supported the lower re-sponse limits used for PS elements. The addition of supplementalnon-PS reinforcement in PCS3b, however, improved the ductility tothat of the non-PS case.

Experimental Program for Dynamic Tests

Four panel configurations were examined under blast loads toassess the accuracy of using the static resistance functions forprediction of the dynamic performance. Table 5 and Fig. 10 presentthe configurations examined under dynamic loading. Two wallswere prestressed with seven-wire strand, and two walls were rein-forced with conventional rebar. M1 and M2 were fabricated fromType 3 cement, while M3 and M4 were made from Type 1 cement.The concrete compressive strength was measured in accordancewith ASTM C39 and is summarized in Table 5. WWR conformedto ASTM A185/A497 (ASTM 2007b, 2006b), and all deformedreinforcement conformed to ASTM A615 (ASTM 2012) orA706 Grade 60 (420 MPa) (ASTM 2009).

The dynamic series consisted of four full-scale detonations ofhigh explosives in the far field. Each wall type was examined intwo experiments that included two detonations each. The first det-onation consisted of a small charge chosen to examine the elasticbehavior of the panels. This detonation is referred to as the prelimi-nary detonation or predetonation. After the predetonation, the pan-els were examined to ensure that no inelastic damage had occurred.These same sets of panels were then subject to a primary detona-tion. The primary detonation consisted of a larger size demandthat was expected to cause significant yielding and damage to thewall systems. The walls were supported by a reinforced concretereaction structure, as illustrated in Figs. 11 and 12. The walls weresupported at the foundation and roof, creating a one-way response.The walls were separated by a small gap to ensure that the panelsdid not interact. The gaps were covered with light-gauge metalsheeting to minimize pressure leakage into the reaction structure.

The responses of the panels were measured using a series ofreflected pressure transducers, load cells, and displacement trans-ducers (Fig. 11). All panels were installed with two load cells inline with the top of the panel to measure the forces generated duringinbound and rebound response. Fig. 12 presents a schematic ofthe setup. The base connection was varied between experiments 1

Fig. 5. Resistance backbone development

Table 2. Average Resistance Backbone Response of Panels

Panel ID Pmax [kPa (psi)]

Measured rotation limits(degrees)

θL θM θS θT

PCS7 (non-PS) 32.0 (4.64) 1.1 5.8 7.0 8.3TS1 (non-PS) 19.3 (2.80) 0.4 5.5 7.0 9.6TS2 (non-PS) 29.0 (4.21) 1.0 5.0 6.7 9.3PCS3a (non-PS) 36.6 (5.31) 0.6 3.9 5.3 6.6PCS4 (PS) 32.1 (4.66) 1.7 5.9 6.2 9.9PCS5 (PS) 33.4 (4.85) 0.7 4.9 6.5 7.3PCS3b (PS) 39.8 (5.77) 0.9 4.6 5.6 5.6PCS8 (PS) 28.7 (4.16) 0.5 4.6 5.8 6.9PCS6 (PS) 29.7 (4.30) 0.3 2.7 4.5 5.3PCS9 (PS) 38.6 (5.60) 0.5 2.0 3.8 4.3PCS2 (PS) 35.2 (5.11) 0.2 1.8 3.0 4.0PCS1 (PS) 30.8 (4.47) 0.4 1.7 4.8 6.9

Fig. 6. Non-prestressed resistance backbones

Fig. 7. Prestressed resistance backbones



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and 2. In experiment 1, two steel restraint angles were used at theexterior of each panel to prevent rebound. This restraint was suc-cessful but produced undesired fixity at the base. Consequently, forexperiment 2, the base restraint was moved to the inside of the wall.This configuration allowed for rotation of the panel and providedrestraint against rebound translation of the base.

Experimental Results of Dynamic Tests

Measured Panel Demands

The pressure demand was measured with three reflected pres-sure gauges located on the reaction structure. Due to detonation

characteristics, the pressures varied across the face of the teststructure. The reflected pressure at the center of each panel waslinearly extrapolated from the three pressure gauge measurements.The maximum and minimum pressure and positive impulsedemand for each panel are summarized in Table 6. The reflectedpressure demand for M1 is illustrated in Fig. 13; the other paneldemands were similar. Demands were increased between experi-ments 1 and 2.

Measured Panel Response

The panels reached their peak inbound deformations after the com-pletion of the positive-pressure phase. In all cases, duration of the

Table 3. Allowable and Measured Response Limits

Component damage level Superficial Moderate Heavy Hazardous

Non-PS flexural limit (USACE 2008) μ ≤ 1.0 2.0° 5.0° 10.0°Non-PS panel response (AVE� SD) 0.78°� 0.34° 5.06°� 1.51° 6.51°� 1.30° 8.44°� 1.97°PS flexural limit (USACE 2008) μ ≤ 1.0 1.0° 2.0° 3.0°PS panel response (AVE� SD) 0.65°� 0.53° 3.48°� 1.60° 5.00°� 1.28° 6.26°� 2.06°

Table 4. Results Summary

Panel ID

Measurements Estimations

Max pressure(psi)

Correspondingdisplacement (in.)

Corresponding momentcapacity (kip-in.)

Estimated foldingcapacity (kip-in.)

Estimated compositecapacity (tension steel)

(kip-in.)

Estimatednoncomposite

capacity (kip-in.)%

Composite

Non-PSTS2 4.2� 0.1 5.3� 0.5 121� 2 73.3 82.2 36.2 185PCS3a 5.3� 1.0 4.1� 2.8 306� 56 198.6 223.7 89.7 162PCS7 4.7� 0.5 6.2� 1.6 267� 31 198.7 263.1 92.8 103TS1 3.1� 0.5 5.0� 1.3 89� 14 84.2 102.8 45.4 77

PSPCS2 5.1� 0.2 1.9� 0.5 295� 10 243.2 272.2 103.6 113PCS9 5.6� 0.2 2.1� 0.4 323� 13 256.6 328.7 116.7 97PCS5 4.9� 0.3 5.1� 0.4 280� 17 231.3 311.4 95.1 89PCS1 4.5� 0.2 1.9� 0.2 259� 11 247.4 227.9 112.5 85PCS4 4.7� 0.4 6.2� 0.8 268� 23 241.6 314.7 105.1 78PCS8 4.0� 0.3 4.1� 1.3 232� 20 227.6 312.4 91.6 63PCS6 4.3� 0.8 2.9� 0.8 250� 45 252.1 323.1 112.2 65PCS3b 5.8� 0.4 4.9� 0.9 332� 21 405.1 489.8 182.8 49PCS3b 5.8� 0.4 4.9� 0.9 332� 21 405.1 489.8 182.8 49PCS3b 5.8� 0.4 4.9� 0.9 332� 21 405.1 489.8 182.8 49

Fig. 8. Strength estimates for sandwich panelFig. 9. Comparative static response



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Table 5. Dynamic Test Matrix

Panel ID Wythe configuration InsulationReinforcement

(longitudinal / transverse) Shear tie type Experimentf 0c interior wytheMPa (ksi)

f 0c exterior wythe,

MPa (ksi)

PSM1 3-2-3 EPS 3=8∅ strand/WWR C 1 60.1 (8.7) 60.1 (8.7)

2 58.7 (8.5) 58.7 (8.5)M2 3-2-3 XPS 3=8∅ strand/#3 A 1 52.7 (7.6) 52.7 (7.6)

2 53.5 (7.8) 53.5 (7.8)Non-PS

M3 3-2-3 XPS #4/#3 A 1 35.1 (5.1) 33.3 (4.8)2 38.4 (5.6) 31.2 (5.2)

M4 6-2-3 XPS #4/#3 and WWR B 1 35.1 (5.1) 33.3 (4.8)2 38.4 (5.6) 31.2 (5.2)

Fig. 10. Specimen sections for dynamic tests



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positive-pressure phase was less than 50% of the wall’s shortestnatural period. At this duration, the demand can be modeled wellas a dynamic loading. The midspan response of each panel is illus-trated for each detonation. Included in each plot is the measuredpressure at the center gauge, MRP2.

Predetonation deformations were limited to a max of 0.22° sup-port rotation, primary detonation 1 to a max of 4.8°, and primarydetonation 2 to a max of 6.8°. Using the USACE (2008) limits, thepanels performed in the superficial damage level for the predeto-nations. Under the primary detonations, the non-PS panels per-formed in the hazardous level, whereas the PS panels, M1 and M2,exceeded the current hazardous limit by achieving 5.7° and 5.6°without failure, respectively. M1, which matched the details usedin specimen PCS2, exceeded the deformation limit of 4.0° mea-sured statically. This indicates that additional performance maybe achieved by the panels at dynamic rates.

Comparisons of the individual responses of the panels werenot consistent due to the variation in impulse imparted to each of

the panels. In general, the panels exhibited elastic response in thepredetonations and inelastic response in the primary detonations.This was notable in the larger periods and permanent deformationobserved in Fig. 14. Due to the variation in demands applied toeach panel, comparisons of the responses were made with respectto analytical predictions using single-degree-of-freedom (SDOF)models.

Estimation of Dynamic Pressure-DisplacementResponse

The dynamic responses of the panels may be estimated using anSDOF approach. The SDOF method simplifies the distributed massand distributed flexural deformation into a single spring-masssystem. This method is commonly used for evaluation of structuralcomponents subjected to blast. The methodology accounts for theelastic and inelastic responses of the wall element and the reflectedpressure demands applied to the walls. The equivalent dampeddynamic equation of motion used is shown as Eq. (4):

KLMðyÞ · M · yðtÞ þ C · _yðtÞ þ RðyÞ ¼ FðtÞ ð4Þ

The SDOF method utilizes an equivalent mass, Me; an appliedforce, FðtÞ; and a component resistance function, RðyÞ, which aredependent on the assumed mode shape, boundary conditions, andexternally applied loading of the actual component. These aspectsof the structural component are accounted for with a load-massfactor, KLM . Damping, C, varies based on the materials usedand changes as damage accumulates in the structural component.Damping was assumed to be 2% for the study. This represents thelower-bound damping value recommended for lightly crackedconcrete components (Oswald and Naito 2012). The dynamicevaluation is solved incrementally using numerical integration.Further details on the SDOF method used can be found inBiggs (1964).

The mass, damping, and demand can be readily deter-mined based on the characteristics of the panel and demand.

Fig. 11. Dynamic reaction structure and setup

Fig. 12. Dynamic test setup



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The resistance function, however, is dependent on proper estima-tion of the panel nonlinearity. Two approaches were examined foraccuracy. In Method 1, the resistance function was determineddirectly from the preceding static tests conducted in this paper.The measured resistance functions from PCS2, TS2, and TS1 wereadopted and applied to M1, M3, and M4, respectively. An equiv-alent static test panel was not available for M2 and therefore wasnot included in the Method 1 comparison. In Method 2, the resis-tance function was determined by estimating the static response ofthe panel from the geometric and material properties. This informa-tion included all dimensions and configurations, concrete compres-sive strength, reinforcement tensile strength, and shear tie shearstrength. The strength and deformations for the resistance functionswere computed as discussed previously using the unmodified UFCand the modified UFC approaches.

The measured resistances of three of the panels were comparedwith the estimated composite and noncomposite strengths. The

noncomposite strength was estimated assuming that each wytheacts independently. Composite performance was estimated withtwo methods. The first estimate was based on the method standard-ized in Unified Facility Criteria 3-340-02 (DoD 2008). This methodcomputes the flexural strength using traditional rectangular stressblock assumptions and ignores the contribution of the steel locatedin the compression face. The approach is consistent with that ofthe PCI Committee on Sandwich Wall Panels (PCI 2011). Dynamicand static increase factors (DIF and SIF) on materials are used toaccount for strength increase at dynamic rates and inherent over-strengthening. Concrete and steel were given SIFs of 1.0 and 1.1and DIFs of 1.19 and 1.17, respectively, in accordance with DoDrecommendations (DoD 2008). The elastic deflections were basedon an initial stiffness computed from the average of the gross andcracked moment of inertias. The second composite estimate isbased on a modification of the UFC approach. The resistance wascomputed using a lower stiffness (i.e., one-quarter of the sum of thegross and cracked moment of inertia) without the addition of staticor dynamic increase factors. Fig. 15 compares the two compositestrength estimates with the experimental results of panels PCS2,TS2, and TS1. Panels PCS2 and TS2 were designed as composite,while TS1 was designed as noncomposite.

The composite and noncomposite response estimates boundthe measured performance. The composite estimate based on UFCmethods provides an overprediction of initial stiffness and strength.This overestimation of the static performance may result in anunder prediction of dynamic deflections and reactions.

The measured dynamic reactions were also compared to esti-mated SDOF reactions to assess the accuracy of the approach.The magnitude of the reaction is dependent on the applied loadand inertial force of the wall. The inertial resistance of the wallis dependent on the boundary conditions of the system, the distri-bution of mass, and the type of loading applied. The dynamic shearreaction, VðtÞ, can be computed as a function of the applied blastdemand, FðtÞ, and the resistance of the panel, RðtÞ. For a uniformmass and load distribution, simple-simple (S-S) boundary condi-tions provide the lowest shear reaction [Eq. (5)], while the fixedsupport of a simple-fixed (S-F) boundary provides the largestreaction [Eq. (6)]. The derivations of these formulations are wellestablished (DoD 2008):

VSimple−SimpleðtÞ ¼ 0.39RðtÞ þ 0.11FðtÞ ð5Þ

VFixed−SimpleðtÞ ¼ 0.43RðtÞ þ 0.19FðtÞ ð6Þ

The measured peak deflection and dynamic reactions are com-pared with the SDOF approaches in Tables 7 and 8. To bound theresponse, the deflections were conservatively estimated assumingan S-S boundary condition. The reactions were estimated using S-Sand the fixed end of an S-F boundary condition. The reactions werealso computed using the equivalent static reaction (ESR). The ESRis equivalent to the reaction force obtained when the panel is at itsstatic flexural capacity, as determined using the modified UFCresistance.

The peak deformation is best estimated using the backboneresistance function of method 1 followed by the modified UFCand standard UFC approaches. To provide a conservative estimateof deformation, the predictive method used should overpredictthe deformation. As illustrated in Table 7, the standard UFC ap-proach on average predicts only 82% of the actual deformation.The modified UFC approach and the backbone methods provide115% and 143% of the measured deformation, respectively. Atlower blast demands, the level of conservatism for the backboneapproach is greater, ranging from 138 to 215% of that measured.

Table 6. Pressure and Positive Impulse Demand

Test Panel ID

Normalized pressure NormalizedpositiveimpulseMaximum Minimum

Experiment 1Pre M1 0.06 −0.021 0.08Primary 0.60 −0.071 0.68Pre M2 0.07 −0.025 0.08Primary 0.72 −0.067 0.79Pre M3 0.07 −0.023 0.08Primary 0.77 −0.067 0.81Pre M4 0.08 −0.018 0.08Primary 0.78 −0.059 0.76

Experiment 2Pre M1 0.10 −0.023 0.09Primary 1.00 −0.070 0.98Pre M2 0.10 −0.022 0.10Primary 0.96 −0.076 1.00Pre M3 0.09 −0.020 0.09Primary 0.88 −0.076 0.94Pre M4 0.08 −0.016 0.08Primary 0.81 −0.080 0.80

Note: Pressure and impulse are normalized by the largest measuredresponse. Dimensional demands can be found in Naito et al. (2011b).

Fig. 13. Normalized pressure demands-Panel M1



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The backbone approach provides the greatest conservatism. Useof this method, however, requires an experimentally generatedresistance function. On average, the modified UFC approachprovides the most accurate estimate of the dynamic deformation,115% of the measured value. Based on the conservatism of the

estimate and the accuracy, this method appears to be best suitedfor design.

Reaction forces are conservatively estimated using an S-Fboundary condition assumption. Use of S-S boundary conditionsprovides a nonconservative estimate of reaction forces for the

Fig. 14. Measured dynamic deformations

Fig. 15. Experimental and estimated resistance functions for three panels



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majority of cases. On average, the assumption of S-S conditionsresults in only 68.3% of the measured reaction force. The as-sumption of an S-F boundary condition provides a conservativeestimate of reaction force for all primary detonations. The smaller-scale predetonation reactions are conservatively predicted by thestandard UFC resistance functions. Use of the modified UFC andbackbone resistance functions results in a underprediction of reac-tions for the predetonations. At this low-level demand, however, theequivalent static reaction is greater than the estimated dynamicreactions. For this case, the equivalent static reaction would controlthe design. On average, the modified UFC method provides the

most accurate estimate of the dynamic reaction, 99.2% of the mea-sured value.

Dynamic Reactions

The measured reactions should be modeled as dynamic loads.The durations of the inbound reactions were on the order of 10to 25 ms in duration, as illustrated in Fig. 16. Fundamental periodsfor the panel range from 40 to 80 ms. The ratio of load duration tocomponent period ranged from 0.12 to 0.63. At these ratios, thedemand should be modeled as a dynamic loading. Consequently,equating the dynamic reactions to equivalent static loads willresult in unrealistically conservative connection designs.

Rebound reactions are dependent on the level of damageincurred by the structural component. For the large demands, min-imal negative dynamic reactions occurred. On average, the negative

Table 7. Comparison of Measured and Estimated Peak Deflections

Panel Method

Primarydetonation Predetonation

1 2 1 2

PSM1 Measured (mm) 87.1 14.9 0.23 0.28

UFC (%) 80 85 102 100UFC Mod. (%) 82 86 153 142Backbone (%) 103 103 178 162

M2 Measured (mm) 75.2 14.5 0.36 0.36UFC (%) 129 95 69 83

UFC Mod. (%) 128 94 104 118Backbone (%) NA NA NA NA

Non-PSM3 Measured (mm) 123.2 17.8 0.56 0.43

UFC (%) 110 89 51 74UFC Mod. (%) 125 100 73 102Backbone (%) 122 93 150 215

M4 Measured 88.1 11.5 0.30 0.25UFC (%) 79 60 45 53

UFC Mod. (%) 139 99 138 164Backbone (%) 126 91 172 200

Table 8. Comparison of Measured and Estimated Reactions

Panel Method

Primary detonation Predetonation

1 2 1 2

S-S S-F(F) S-S S-F(F) S-S S-F(F) S-S S-F(F)

PSM1 Measured 203.97 kN (45.9 kip) 448.4 kN (100.8 kip) 67.22 kN (15.1 kip) 94.33 kN (21.2 kip)

E.S.R. (%) 45.1 62.6 20.5 28.5 136.9 189.8 97.6 135.3UFC (%) 95.0 152.2 80.9 129.6 82.6 135.9 69.4 103.7

UFC Mod. (%) 86.8 150.5 73.0 123.8 60.4 67.0 44.1 52.4Backbone (%) 86.8 149.8 71.0 119.1 58.5 65.0 42.8 52.3

M2 Measured 322.0 kN (72.4 kip) 360.6 kN (81.1 kip) 82.38 kN (18.5 kip) 66.6 kN (15.0 kip)E.S.R. (%) 27.2 37.6 24.2 33.6 106.2 147.1 131.3 181.9UFC (%) 66.2 114.6 103.3 161.8 69.3 108.9 99.4 144.5

UFC Mod. (%) 66.0 114.1 87.4 148.2 51.4 57.4 64.4 74.1Backbone (%) NA NA NA NA NA NA NA NA

Non-PSM3 Measured 292.7 kN (65.8 kip) 425.6 kN (95.7 kip) 62.2 kN (14.0 kip) 82.7 kN (18.6 kip)

E.S.R. (%) 15.8 29.6 10.9 20.4 74.3 139.3 55.9 104.8UFC (%) 77.9 134.5 76.8 112.2 74.2 114.5 57.4 91.2

UFC Mod. (%) 77.7 134.2 66.1 116.1 55.8 84.9 43.1 66.4Backbone (%) 77.3 133.5 61.2 105.2 34.8 60.1 32.8 56.4

M4 Measured 338.9 kN (76.2 kip) 286.3 kN (64.4 kip) 57.9 kN (13.0 kip) 65.0 kN (14.6 kip)E.S.R. (%) 7.3 11.6 8.6 13.7 42.7 68.0 38.1 60.5UFC (%) 68.5 118.8 85.3 149.9 103.3 126.0 91.1 111.7

UFC Mod. (%) 68.0 117.6 83.6 144.4 42.5 72.9 37.0 63.9Backbone (%) 68.0 117.3 83.6 144.3 42.2 72.5 37.0 63.8

Fig. 16. Measured reaction demands



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reaction demand was 20% of the inbound for primary detonations1 and 2. For small demands generated from preliminary detonations1 and 2, the rebound reactions were 95% of the inbound reactionson average.

Conclusions

An experimental study was conducted to assess the performancesof non-load-bearing insulated concrete sandwich wall panels sub-jected to demands generated from far-field detonation of highexplosives. The study examined the resistance provided under staticuniform loads as well as full-scale blast demands. The measuredstatic and dynamic deformations and reactions were compared withcommon predictive methods to assess their applicability for mod-eling PS and non-PS concrete wall panels. Based on the resultspresented, the following conclusions were made:1. In general, the PS panels were not able to reach their estimated

composite flexural strength, whereas the majority of non-PSconcrete panels did achieve their composite capacity.

2. The ductility of the PS panels was less than that of the non-PSpanels. The use of prestressing alone resulted in a 50% reduc-tion in ductility from a similar sandwich panel design utilizingnon-PS reinforcement. The combined use of PS and non-PSreinforcement, however, resulted in ductility levels similarto that of a non-PS panel.

3. The postcracking performance of the panels was sensitiveto the tie and foam types used. Stiff, brittle ties resulted ina stiff, brittle flexural response, and a flexible ductile tie suchas a C-clip provided a flexible ductile flexural response.The use of EPS and Polyiso insulation provided more shearresistance and composite action than the use of XPS foam.This was attributed to the rough, absorbent surface of theEPS and Polyiso, which increased the shear transfer betweenwythes.

4. Due to the shear tie slip and fracture, and the compressibilityof the insulation, a distributed flexural hinge did not occurunder static loading; instead the systems exhibited a foldingmechanism. A simplified approach was used to predict thisreduced flexural capacity and was found to provide a conser-vative estimate of the static strength.

5. The measured static deformation capacities of the panels metall USACE response limits for prestressed concrete flexuralelements. The non-PS panels met the USACE response limitsup to a heavy level of damage. Due to the folding mechanismobserved, the non-PS panels were not capable of achieving 10°of support rotation.

6. Under static demands, both non-PS and PS concrete sandwichpanels exhibited an elastic-hardening load deformation re-sponse up to the peak capacity. After the maximum capacitywas achieved, the systems exhibited an extended softeningresponse. This behavior deviated from the elastic-plasticperformance expected using UFC methods.

7. The insulated PS panel subjected to blast loading exceededthe USACE response limits for hazardous damage withoutsignificant damage. The non-PS panels performed in thehazardous range without significant observed damage. ThePS panels supported dynamic displacements in excess of theirmeasured static displacement capacity. This indicates that per-formance enhancements occurred at dynamic rates. It is likelythat the insulation foam may have provided enhanced shearresistance and improvements in composite action due tothe rate sensitivity of the material. This should be examinedfurther.

8. The addition of supplemental non-PS reinforcement to a PSpanel improved the ductility to that of the non-PS case.Therefore, for panels for which large rotations are needed,addition of non-PS reinforcement should increase the ductilityof the system.

9. The dynamic responses of three of the panels were estimatedusing simplified SDOF analyses. A resistance function wasdetermined using a modified UFC method in which the initialstiffness was based on one-fourth of the cracked and grosssection inertia with the assumption of no dynamic increasefactor for concrete or reinforcement. This method was foundto provide an accurate and conservative means of estimatingthe peak inbound deformations.

10. Measured reactions were found to be of short duration andshould be modeled as dynamic loads. The reactions were bestestimated using the modified UFC method and a standardSDOF approach with simple-fixed boundary conditions. Useof S-S boundary conditions will provide a nonconservativeestimate of the support demands.

Acknowledgments

The authors thank the Air Force Research Laboratory (Dr. RobertDinan and Dr. Michael Hammons, Program Managers) for fund-ing this work under contracts FA4918-07-D-0001 and FA8903-08-D-8768. The experiments were performed at the Air ForceResearch Laboratory located at Tyndall AFB, FL, and at theUniversity of Missouri – Columbia. The work was conducted undera Cooperative Research and Development Agreement (CRADA)between the Portland Cement Association, the Prestressed/PrecastConcrete Institute, and the Tilt-Up Concrete Association. Theauthors also thank John Sullivan, Jason Krohn, Michael Sugrue,and the member companies of these organizations for technicaland fabrication support of the research. In addition, the authorsthank the Geotechnical and Structures Laboratory of the U.S. ArmyCorps of Engineers for efforts in revising this document. Citationof manufacturers or trade names does not constitute an officialendorsement or approval of the use thereof. The U.S. Governmentis authorized to reproduce and distribute reprints for Governmentpurposes notwithstanding any copyright notation hereon.

References

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ASTM. (2005). “Standard test method for compressive strength ofcylindrical concrete specimens.” C39, West Conshohocken, PA.

ASTM. (2006a). “Standard specification for steel strand, uncoatedseven-wire for prestressed concrete.” A416, West Conshohocken,PA.

ASTM. (2006b). “Standard specification for steel welded wirereinforcement, deformed, for concrete.” A497, West Conshohocken,PA.

ASTM. (2007a). “Standard specification for steel wire, plain, for concretereinforcement.” A82, West Conshohocken, PA.

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