July 2015 

NASA/TM–2015-218772 

Multi-Level Experimental and Analytical Evaluation 

of Two Composite Energy Absorbers 

Karen E. Jackson and Justin D. Littell 

Langley Research Center, Hampton, Virginia 

Edwin L. Fasanella 

National Institute of Aerospace, Hampton, Virginia 

Martin S. Annett 

Langley Research Center, Hampton, Virginia 

Michael D. Seal, II 

Analytical Mechanics Associates, Inc., Hampton, Virginia 
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July 2015 

NASA/TM–2015-218772 

Multi-Level Experimental and Analytical Evaluation 

of Two Composite Energy Absorbers 

Karen E. Jackson and Justin D. Littell 

Langley Research Center, Hampton, Virginia 

Edwin L. Fasanella 

National Institute of Aerospace, Hampton, Virginia 

Martin S. Annett 

Langley Research Center, Hampton, Virginia 

Michael D. Seal, II

Analytical Mechanics Associates, Inc., Hampton, Virginia
Available from: 

NASA STI Program / Mail Stop 148 

NASA Langley Research Center 

Hampton, VA 23681-2199

Fax: 757-864-6500

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Space Administration. 
1 

Multi-Level Experimental and Analytical Evaluation of Two Composite Energy Absorbers 

 Karen E. Jackson Justin D. Littell Edwin L. Fasanella 

NASA Langley Research Center NASA Langley Research Center National Institute of Aerospace

 Martin S. Annett Michael D. Seal 

 NASA Langley Research Center Analytical Mechanics Associates, Inc. 

Mail Stop 495 

NASA Langley Research Center 

Hampton, VA 23681 

1.0 ABSTRACT 

Two composite energy absorbers were developed and evaluated at NASA Langley Research Center 

through multi-level testing and simulation performed under the Transport Rotorcraft Airframe 

Crash Testbed (TRACT) research program. A conical-shaped energy absorber, designated the 

conusoid, was evaluated that consisted of four layers of hybrid carbon-Kevlar® plain weave fabric 

oriented at [+45°/-45°/-45°/+45°] with respect to the vertical, or crush, direction. A sinusoidalshaped energy absorber, designated the sinusoid, was developed that consisted of hybrid carbonKevlar® plain weave fabric face sheets, two layers for each face sheet oriented at ±45° with respect 

to the vertical direction and a closed-cell ELFOAM® P200 polyisocyanurate (2.0-lb/ft3

) foam core. 

The design goal for the energy absorbers was to achieve average floor-level accelerations of 

between 25- and 40-g during the full-scale crash test of a retrofitted CH-46E helicopter airframe, 

designated TRACT 2. Variations in both designs were assessed through dynamic crush testing of 

component specimens. Once the designs were finalized, subfloor beams of each configuration were 

fabricated and retrofitted into a barrel section of a CH-46E helicopter. A vertical drop test of the 

barrel section was conducted onto concrete to evaluate the performance of the energy absorbers 

prior to retrofit into TRACT 2. The retrofitted airframe was crash tested under combined forward 

and vertical velocity conditions onto soil, which is characterized as a sand/clay mixture. Finite 

element models were developed of all test articles and simulations were performed using LSDYNA®, a commercial nonlinear explicit transient dynamic finite element code. Test-analysis 

results are presented for each energy absorber as comparisons of time-history responses, as well as 

predicted and experimental structural deformations and progressive damage under impact loading 

for each evaluation level. 

2.0 INTRODUCTION 

In 2012, the NASA Rotary Wing (RW) Crashworthiness Program [1] initiated the Transport 

Rotorcraft Airframe Crash Testbed (TRACT) research program by obtaining two CH-46E 

helicopter airframes from the Navy CH-46E Program Office (PMA-226) at the Navy Flight 

Readiness Center in Cherry Point, North Carolina. Full-scale crash tests were planned to assess 

dynamic responses of transport-category rotorcraft under combined forward and vertical impact 

loading. The first crash test, TRACT 1 [2], was performed at NASA Langley Research Center’s 

Landing and Impact Research Facility (LandIR). Impact tests conducted at LandIR provide data that 

enable the study of critical interactions between the airframe, seat, and occupant during a controlled 

crash environment. The CH-46E airframe is categorized as a medium-lift rotorcraft with length and 

width of 45- and 7-ft, respectively, and a capacity for 5 crew and 25 troops. TRACT 1 was 
2 

conducted in August 2013 under combined conditions of 300-in/s (25-ft/s) vertical and 396-in/s (33-

ft/s) forward velocity onto soil, which is characterized as a sand/clay mixture. The primary 

objectives for TRACT 1 were to assess improvements in occupant loads and flail envelope with the 

use of crashworthy features such as pre-tensioning active restraints and energy absorbing seats and 

to develop novel techniques for photogrammetric data acquisition to measure occupant and airframe 

kinematics. Pre- and post-test photographs of the TRACT 1 crash test are shown in Figure 1. 

Figure 1. Pre- and post-test photographs of the TRACT 1 full-scale crash test. 

The TRACT 1 airframe was tested in a baseline configuration with no changes to the structural 

configuration, including the discrete aluminum shear panels in the subfloor. It is important to note 

that the CH-46E does not contain a center keel beam; hence the airframe relies on the aluminum 

shear panels, the cargo rails in the floor, and the airframe structure to provide longitudinal and 

torsional stiffness. A final objective of TRACT 1 was to generate crash test data in a baseline 

configuration for comparison with data obtained from a similar TRACT 2 crash test. The crash test 

of the second CH-46E airframe (TRACT 2) was conducted on October 1, 2014 and was performed 

for the same nominal impact velocity conditions and onto the same sand/clay surface [3]. The 

difference is that the TRACT 2 airframe was retrofitted with three different composite energy 

absorbing subfloor concepts located in the mid-cabin region: a corrugated web design [4, 5] 

fabricated of graphite fabric; a conical-shaped design, designated the “conusoid,” fabricated of four 

layers of hybrid carbon-Kevlar® fabric [6]; and, a sinusoidal-shaped foam sandwich design, 

designated the “sinusoid,” fabricated of the same hybrid fabric face sheets with a foam core. While 

the TRACT 2 airframe contained similar seat, occupant, and restraint experiments, one of the major 

goals of the test was to evaluate the performance of novel composite energy absorbing subfloor 

designs for improved crashworthiness. 

This paper will summarize the development of the conusoid and sinusoid foam sandwich energy 

absorbing concepts. Multi-level evaluations of the energy absorbers are discussed including 

dynamic crush testing and simulation of component specimens, vertical drop testing and simulation 

of a retrofitted barrel section, and full-scale crash testing and simulation of the TRACT 2 retrofitted 

helicopter airframe. Finite element models were developed of all test articles and simulations were 

performed using LS-DYNA® [7, 8], a commercial explicit nonlinear, transient dynamic finite 

element code. Thus, a final objective of this research program is to evaluate the capabilities of LSDYNA® in predicting the dynamic response and progressive failure behavior of composite energy 

absorbing airframe structures. 
3 

3.0 DESCRIPTION OF THE COMPOSITE ENERGY ABSORBERS 

3.1 Design Goals for the Energy Absorbers 

Following the TRACT 1 crash test, a research effort was initiated to develop two composite energy 

absorbers for retrofit into the TRACT 2 test article. The design goals were to limit the average 

vertical accelerations to 25- to 40-g on the floor, to minimize peak crush loads, and to generate 

relatively long crush stroke limits under dynamic loading conditions, typical of those experienced 

during the TRACT 1 full-scale crash test [2]. To further clarify the design goals, it is important to 

note that the loading conditions on the frames of the TRACT 1 full-scale crash test provided 

dynamic crush loads of approximately 2,500- to 4,000-lb. per linear foot, measured from one side of 

the floor to the other (a distance of 60-in. or 5-ft). These values are determined by multiplying the 

design acceleration levels (25- to 40-g) by the floor mass loading of 100-lb per linear foot. Note 

that the weight times the g-factor equals the force. The loading condition was based on seat and 

occupant loads that were recorded during the TRACT1 crash test. A schematic drawing is shown in 

Figure 2 illustrating design conditions for floor loading. 

Figure 2. Floor loading condition schematic. 

Figure 2 shows an idealized schematic of the floor and subfloor located at an individual fuselage 

frame. The floor, which is approximately 5-ft wide, is divided into 5 segments of 1-ft. length, each 

having an associated floor loading of 100-lb. The energy absorbers, depicted as individual springs, 

are designed to limit floor-level accelerations to 25- to 40-g. 

Often, design goals for energy absorbers are defined in terms of Specific Sustained Crush Stress 

(SSCS). The SSCS is a measure of the energy absorbing capability of the material and is defined as 

 SSCS = Pavg/(A x ρ) Eq. 1 

where Pavg is the average sustained crush load, A is the cross-sectional area, and ρ is the density of 

the material. The SSCS is also the energy absorbed per unit weight of material crushed. Assuming 

an average acceleration level of 25-g, a floor loading of 100 lb. for a 1-ft. length, a cross-sectional 

area of 0.98-in2 and a material density of 0.03486-lb/in3, an SSCS value of 72,590-in-lb/in is 

obtained for a typical conusoid energy absorber. This value is particularly high. For example, 

Reference 9 documents the energy absorption capabilities of flat plate composite specimens and 

reports values of SSCS ranging from 28,710- to 40,200-in-lb/in. Part of the explanation for the high 

SSCS value is the fact that an average crush load is based on the dynamic design goal. In Reference 

6, the average crush load for the conusoid is reported between 900- to 1,500-lb, based on quasistatic loading. Using these averages, SSCS is lowered to values between 26,132- and 43,611-in-
4 

lb/in. The SSCS is typically reported in metric units. Thus, a SSCS value of 72,590-in-lb/in is 

converted to 18.1-N m/g. Note that Farley [10] reported SSCS values of between 20- and 75-Nm/g 

for various composite tubes that were subjected to quasi-static compressive loads.

Farley [10, 11], Kindervater [12], Bannerman [13], and Hanagud [14] have investigated the 

crushing response of composite structural elements and sine wave beams. Farley [11] has shown 

that high values of SSCS are obtained when using hybrid graphite-Kevlar composites in which the 

graphite fibers are oriented in the same direction as the loading axis and the Kevlar fibers are 

oriented at 45° to the loading axis. As stated in Reference 11, “the Kevlar fibers are positioned in 

the laminate to provide containment and support for the graphite fibers, which absorb energy 

through a combination of crushing and fracturing modes.” 

3.2 Conusoid Energy Absorber 

The first design is a novel conical sinusoid, or “conusoidal” composite energy absorber, also 

designated “conusoid.” The geometry of the conusoid is based on alternating right-side-up and upside down half-cones placed in a repeating pattern. The conusoid combines a simple cone design, 

which has been extensively studied in the literature [15-18], with sinusoidal beam geometry to 

create a structure that utilizes the advantages of both configurations. An isometric view of the 

conusoid is shown in Figure 3(a). Variations in geometry, materials, and laminate stacking 

sequences were evaluated during development of the conusoid and the final design consisted of four 

layers of a hybrid carbon-Kevlar® plain weave fabric oriented at [+45°/-45°/-45°/+45°] with respect 

to the vertical, or crush, direction. A photograph of a typical conusoid component is shown in 

Figure 3(b). Dimensions of the component are 12-in. long, 7.5-in. high, with an overall width of 

1.5-in. Additional information on the development and fabrication of the conusoid energy absorber 

may be found in Reference 6. 

 (a) Schematic drawing. (b) Photograph of component. 

Figure 3. Isometric view and photograph of a conusoid component. 

3.3 Sinusoid Foam Sandwich Energy Absorber 

The second energy absorber, designated the “sinusoid,” consisted of hybrid carbon-Kevlar® plain 

weave fabric face sheets, two layers for each face sheet oriented at ±45° with respect to the vertical, 

or crush, direction and a closed-cell ELFOAM® P200 polyisocyanurate (2.0-lb/ft3) foam core. Sine 

wave energy absorbers have been studied extensively because they offer desirable features under 

compressive loading [19-22]. Energy absorption values from sine wave concepts can be similar to 
5 

values obtained from crush tubes. In addition, sine wave concepts tend to deform in a stable 

manner through plastic hinge formation and crushing, rather than global buckling. Often, the actual 

shape of the energy absorber is not truly a sine wave, but a series of alternating half circles. In fact, 

the sinusoid concept described in this paper is actually a series of half circles with a diameter of 

1.75-in.; however, the designation of “sinusoid” will continue to be used. 

The total thickness of a sinusoid component was 1.5-in. with a length of 12-in. and a height of 7.5-

in. Design parameters were assessed through component testing including different materials for the 

face sheets and different laminate stacking sequences. Variations in sinusoid geometry were not 

evaluated since an existing mold was used in construction. A photograph of a sinusoid foam 

sandwich specimen is shown in Figure 4. Note that, in preparation for the component drop test, 0.5-

in.-thick polycarbonate plates were attached to both the top and bottom surfaces of the specimen. 

 Figure 4. Pre-test photograph of a sinusoid foam sandwich component. 

4.0 COMPONENT TESTING AND SIMULATION 

4.1 Conusoid Energy Absorber 

Different conusoid energy absorber designs were dynamically crushed in a 14-ft. drop tower with an 

instrumented 110-lb. falling mass. The impact condition for all of the dynamically crushed 

specimens was approximately 264-in/s (22-ft/s). The drop mass was instrumented with a 500-g 

damped accelerometer and data were acquired using a National Instruments Data Acquisition 

System (DAS) sampling at 25-kHz. All post-processed acceleration data were filtered using a lowpass 4-pole Butterworth filter with a 500-Hz cutoff frequency. A high-speed camera filming at 1-

kHz captured the deformation time history, which is depicted in Figure 5. The identified failure 

mechanism is folding of the conusoid walls, which is a desirable failure mode that produces a stable 

and constant crush response within the design level of 25-40 g. 

A depiction of the finite element model representing the conusoid energy absorber is shown in 

Figure 6. The model contained 185,940 nodes; 44,294 Belytschko-Tsay shell elements; 116,100 

solid elements representing the rigid drop mass, 1 initial velocity card assigned to nodes forming the 

rigid mass, and 1 body load card defining gravity. The nominal shell element edge length is 0.032-
6 

in. The shell elements representing the hybrid carbon-Kevlar® fabric layers were assigned Mat 58, 

which is a continuum damage mechanics material model used in LS-DYNA® for representing 

composite laminates and fabrics [23]. 

Figure 5. High-speed video clips of conusoid deformation. 

 (a) Model with impact mass. (b) Model without impact mass. 

Figure 6. Depictions of the conusoid component model. 

Baseline Mat 58 properties are listed in Table 1. Properties for Mat 58 were obtained through 

detailed test-analysis comparisons with experimental data obtained from standard material 

characterization tests, such as tensile testing of fabric coupons oriented at 0°, 90°, and ±45° to 

obtain longitudinal stiffness and strength, transverse stiffness and strength, and shear stiffness and 

strength, respectively. Once verified through comparison with coupon test data, these properties 

were unchanged for all subsequent simulations of the energy absorbers. It should be noted that Mat 

58 includes certain parameters, such as the SLIM parameters and ERODS, that cannot be 

determined entirely based on experimental data. For these parameters, estimates were input based 
7 

on past experience of the analysts. For the conusoid, individual ply layers were input using the 

*PART_COMPOSITE feature in LS-DYNA® which allows input of ply orientations, ply 

thicknesses, and ply material designations for each layer within a composite laminate. Single Point 

Constraints (SPCs) were used to constrain the nodes forming the bottom plate. The conusoid model 

was executed using LS-DYNA® SMP version 971 on a Linux-based workstation with 8 processors 

and required 11 hours and 49 minutes of clock time to execute the simulation for 0.035-seconds. 

Model output included time-history responses of the drop mass, and image sequences of structural 

deformation. 

Table 1. Mat 58 material properties used to represent hybrid carbon-Kevlar® fabric. 

Material Property Description Symbol Values 

Density, lb-s2/in4 RO 0.903E-4 

Young’s modulus longitudinal direction, psi EA 6.3E+6 

Young’s modulus transverse direction, psi EB 2.76E+6 

Poisson’s ratio, ν21 PRBA 0.03 

Stress limit of nonlinear portion of shear curve, psi TAU1 4,500. 

Strain limit of nonlinear portion of shear curve, in/in GAMMA1 0.06 

Shear modulus AB, BC, and CA, psi GAB 3.0E+5 

Min stress factor for limit after max stress (fiber tension) SLIMT1 0.8 

Min stress factor for limit after max stress (fiber comp) SLIMC1 1.0 

Min stress factor for limit after max stress (matrix tension) SLIMT2 0.8 

Min stress factor for limit after max stress (matrix comp) SLIMC2 1.0 

Min stress factor for limit after max stress (shear) SLIMS 1.0 

Material axes option (model dependent) AOPT 0.0 

Maximum effective strain for element layer failure ERODS 0.5 

Failure surface type FS -1.0 

Strain at longitudinal compressive strength, in/in E11C 0.007 

Strain at longitudinal tensile strength, in/in E11T 0.0143 

Strain at transverse compressive strength, in/in E22C 0.012 

Strain at transverse tensile strength, in/in E22T 0.025 

Strain at shear strength, in/in GMS 0.45 

Longitudinal compressive strength, psi XC 40,000. 

Longitudinal tensile strength, psi XT 89,000. 

Transverse compressive strength, psi YC 25,000. 

Transverse tensile strength, psi YT 54,000. 

Shear strength, psi SC 7,100. 

Comparisons of predicted and experimental acceleration and displacement time histories of the drop 

mass are shown in Figures 7(a) and (b), respectively. The conusoid model over predicts the 

magnitude of the initial peak acceleration, 96-g compared with 61-g for the test. However, other 

than that anomaly, the level of agreement is good. The average acceleration calculated for the test 

is 28.0-g for pulse duration of 0.0- to 0.025-s, whereas the model average acceleration is 28.4-g for 

the same duration. The results of the conusoid component test indicate that the configuration of the 

energy absorber meets all of the design goals, including achieving a sustained average acceleration 

level of between 25-40-g. The comparison of vertical displacement time histories also exhibits 

good agreement, as shown in Figure 7(b). The maximum displacement of the test article is 2.9-in., 

providing a crush stroke of 38.7%. The maximum displacement of the model is 2.53-in., providing a 
8 

crush stroke of 33.7%. Note that the entire range of crush stroke was not evaluated during 

component testing. 

 (a) Acceleration responses. (b) Displacement responses. 

Figure 7. Acceleration and displacement comparisons for the conusoid component vertical drop test. 

A sequence of model deformation is shown in Figure 8. Stable crushing occurs through plastic 

hinge formation and folding, along with some local buckling of the conusoid walls. The predicted 

response matches the model deformation captured by the high-speed camera, as shown in Figure 5. 

Note that the ERODS parameter is set to 0.5 (see Table 1), indicating that elements are deleted from 

the simulation once they reach an effective strain of 50%. If the erosion parameter is set too low, 

holes produced in the model by element deletion could lead to premature failure and unstable model 

behavior. 

Time = 0.005-s. Time=0.01-s. 

Time=0.015-s. Time=0.02-s. 

Figure 8. Deformation sequence of the conusoid. Note that deleted elements are highly distorted. 
9 

4.2 Sinusoid Energy Absorber 

The sinusoid foam sandwich energy absorber was initially evaluated through quasi-static and 

dynamic crush testing of components. A post-test photograph of a typical sinusoid component is

shown in Figure 9 for a dynamic crush test in which a 113.5-lb mass impacted the sinusoid at 265-

in/s (22.08-ft/s). The sinusoid component is approximately 12-in. long, 7.5-in. high, and 1.5-in. 

wide. A flat 0.5-in.-thick polycarbonate plate was glued to both the top and bottom surfaces of the 

specimen. As shown in Figure 9, the specimen exhibits stable, plastic-like deformation with 

uniform folding of the face sheets and crushing of the foam core. Crushing initiates along the top 

edge of the specimen. Note that the sides of the specimen were not covered with face sheets, which 

allowed splaying of the foam core. 

The LS-DYNA® finite element model representing the sinusoid component drop test is shown in 

Figure 10. The model contained: 53,540 nodes; 7,380 Belytschko-Tsay shell elements; 37,515 solid 

elements; a rigid drop mass; 1 initial velocity card assigned to nodes forming the rigid drop mass; 

SPCs to fully constrain the bottom nodes of the sinusoid; 1 automatic single surface contact; and 3 

material definitions. As with the conusoid, the shell elements were assigned Mat 58, using the 

properties listed in Table 1. The nominal element edge length in the sinusoid model was 0.2-inches. 

 Figure 9. Post-test photograph of a sinusoid foam sandwich energy absorber. 

The solid elements representing the foam core were assigned Mat 63, which is a crushable foam 

material model in LS-DYNA® that allows user input of the stress-strain response of the material in 

tabular format. The stress-strain response of the P200 foam was determined through quasi-static 

testing of 4-in. x 4-in. x 3-in. rectangular blocks. A plot of the experimental curve obtained at a 

crush rate of 1.0-in/minute is shown in Figure 11, along with the stress-strain response used as input 

to Mat 63. Note that the input curve matches the test data to a strain of 0.67-in/in. At this point, the 

test data ends, yet the Mat 63 input response continues and increases dramatically up to 100,000-psi 

at 1-in/in (note that this data point is not shown in the plot). The large “tail” added to the end of the 

stress-strain response represents compaction of the foam and is needed to stabilize the response of 

the solid elements for high values of volumetric strain. 
10 

 (a) Sinusoid component model. (b) Model without drop mass. 

Figure 10. Depictions of the finite element model of the sinusoid component. 

The sinusoid model was executed using LS-DYNA® SMP version 971 on a Linux-based 

workstation with 8 processors and required 10 hours and 34 minutes of clock time to execute the 

simulation for 0.04-seconds. Model output included time-history responses of the drop mass, and 

image sequences of structural deformation. 

Figure 11. Plot of P200 foam stress-strain response. 

Test-analysis comparisons of time-history responses are plotted in Figure 12 for the sinusoid 

component crush test. These results demonstrate excellent test-analysis agreement. As can be seen 

in Figure 12(a), the acceleration response of the drop mass achieves an initial peak of 55-g, then 

drops to approximately 22-g, where it remains constant until the end of the pulse. The model 

mimics this response, even predicting the unloading response near the end of the pulse. The 

average acceleration calculated for the test is 21.8-g for pulse duration of 0.0- to 0.03-s, whereas the 

average acceleration of the predicted response is 22.9-g for the same duration. The experimental 

and analytical displacement responses, shown in Figure 12(b), exhibit maximum values of 4- and 

3.8-in., respectively, which represents approximately 50% stroke. The average acceleration results 

for the sinusoid fall slightly below the required design goal of 25- to 40-g. The lower average crush 

acceleration for the sinusoid translates into a larger crush stroke than was seen for the conusoid. 
11 

 (a) Acceleration responses. (b) Displacement responses. 

Figure 12. Test-analysis time history comparisons for the sinusoid component. 

The predicted sinusoid model deformation is shown in Figure 13 for six discrete time steps. The 

model exhibits stable crushing through folding and plastic-like deformation of the face sheets and 

crushing of the foam core. The deformation pattern matches the test response shown in Figure 9.

 (a) Time=0.005-s. (b) Time=0.01-s. (c) Time=0.015-s. 

 (d) Time=0.02-s. (e) Time=0.025-s. (f) Time=0.03-s. 

Figure 13. Predicted sinusoid model deformation. 

5.0 RETROFITTED BARREL SECTION DROP TESTING AND SIMULATION 

5.1 Experimental Program 

Following the TRACT 1 crash test, a portion of the forward cabin (Fuselage Station FS164 through 

FS250) was separated from the remainder of the post-test specimen. The resulting barrel section 

was essentially undamaged during the TRACT 1 test and was retrofitted with two of the energy 

absorbing concepts planned for TRACT 2, including the conusoid and the sinusoid foam sandwich 

designs. The original floor in the barrel section was removed and was replaced with a sheet of 0.5-

in.-thick polycarbonate. The reason for this change was to enable viewing of the crushing response 

of the energy absorbers using high-speed cameras. The total weight of the fully loaded barrel 
12 

section was 1,810-lb. It was impacted onto concrete at 297.6-in/s (24.8-ft/s). The objectives of the 

barrel section test were to evaluate: (1) the performance of the two energy absorbers during a fullscale drop test prior to the TRACT 2 test, (2) the fabrication methods for the energy absorbers, (3) 

the structural integrity of the retrofit, (4) the strength of the polycarbonate floor, and (5) imaging 

techniques used during the test. 

Pre-test photographs of the barrel section test article are shown in Figures 14(a) and (b), 

highlighting front and rear views, respectively. Close-up photographs showing the conusoid and 

sinusoid foam sandwich energy absorbers that were retrofitted into the barrel section are shown in 

Figures 15(a) and (b), respectively. The conusoid subfloor was located at FS220. The floor of the 

fuselage section was loaded with a 681-lb concentrated mass, which was centered about FS220 and 

located just above the conusoid energy absorber. The sinusoid energy absorber was located at 

FS190, in front of the conusoid. Both energy absorbers were painted white and vertical tick marks 

were added to aid in deformation tracking. 

 

The floor of the fuselage section was loaded with a 320-lb concentrated mass on the right-hand side 

and a double seat with two 50th percentile male Anthropomorphic Test Devices (ATDs) on the lefthand side, both centered about FS190 and located just above the sinusoid energy absorber. The seat 

was attached to the floor using standard seat tracks and the seated dummies were restrained using 

lap belts only. The combined weight of the seat and dummies was 405-lb. This loading condition 

was intended to replicate the anticipated TRACT 2 configuration. 

 (a) Front view. (b) Rear view. 

Figure 14. Front and rear view photographs of the CH-46E barrel section. 

The barrel section test was conducted by raising the test article to a height of 115.2-in. (9.6-ft) and 

releasing it to impact a concrete surface at 297.6-in/s (24.8-ft/s). Data were collected at 25-kHz 

primarily from accelerometers located throughout the cabin and on the 320- and 681-lb masses. 
13 

(a) Photograph of the conusoid energy absorber. 

(b) Photograph of the sinusoid foam sandwich energy absorber. 

Figure 15. Pre-test photographs of the conusoid and sinusoid energy absorbers, as retrofitted 

into the barrel section. 

A post-test photograph highlighting the crushing response of the conusoid energy absorber is shown 

in Figure 16. The original height of the conusoid was 8.31-in., as measured at the center of the 

subfloor, and the post-test height was 3.44-in., providing a crush stroke of 58%. As seen in Figure 

16, the conusoid energy absorber displayed fracturing and delamination of the hybrid carbonKevlar® plies. Most of the damage was located in the center of the specimen, under the 681-lb mass. 

Figure 16. Post-test photograph of the conusoid energy absorber. 

Post-test photographs of the barrel section and the sinusoid energy absorber are highlighted in 

Figure 17. The original height of the sinusoid energy absorber was 8.75-in., as measured at the 

center of the subfloor, and the post-test height was 4.44-in., providing a crush stroke of 49.3%. As 

seen in Figure 17(b), the sinusoid foam sandwich energy absorber displayed crushing of the foam 

core, and fracturing of the face sheets starting from the bottom, curved edge. The sinusoid energy 

absorber in the barrel section did not exhibit the uniform folding pattern seen in the face sheets of 

the component specimen. 

One issue with the barrel section drop test was noted following the impact event. Two 0.5-in.-

diameter steel bolts were used to attach the 320-lb mass to the polycarbonate floor of the section. 

These bolts were approximately 12-in. in length, and one bolt was located 12-in. in front of the 

other. After drilling through the lead mass and floor, the bolts extended approximately 3.5- to 4-in. 

into the subfloor crush zone. The extra bolt lengths were not trimmed prior to the test and the bolts 
14 

impacted the bottom skin of the fuselage section during the test. The forward bolt can be seen 

highlighted on the left side of Figures 15(b) and 17(b). Following the test, the bolts were removed 

and they are depicted in the photograph of Figure 18. The forward bolt impacted the fuselage skin 

and deformed plastically. Much less permanent deformation is seen in the second bolt. The 

presence of these bolts affected the acceleration response of the 320-lb mass during the latter 

portion of the time history response, which will be discussed later in the paper. 

(a) Post-test photograph of the barrel section. 

(b) Post-test photograph highlighting the post-test crush response of the sinusoid. 

Figure 17. Post-test photographs of the barrel section drop test. 

Figure 18. Post-test photograph of exposed steel bolts. 

5.2 Barrel Section Finite Element Model 

The finite element model of the barrel section is shown in Figure 19. This model contains: 105,986 

nodes; 22 parts; 10 material definitions; 57,041 Belytschko-Tsay shell elements; 63,591 solid 

elements; 1,677 beam elements; 1 initial velocity; 1 contact definition; 20 discrete masses 

representing the double seat and ATD occupants; 2 lumped masses representing the 320- and 681-lb 
15 

blocks used in the test article; and 1 planar rigid wall representing the impact surface, which is not 

shown in Figure 19. The seat and occupants were represented using 20 discrete masses assigned to 

nodes at the approximate seat track attachment locations. All nodes in the barrel section model were 

assigned an initial velocity of 297.6-in/s (24.8-ft/s), matching the measured test velocity. The hybrid 

carbon-Kevlar® fabric material, used in the construction of the conusoid and the face sheets of the 

sinusoid foam sandwich energy absorbers, was represented using Mat 58 in LS-DYNA® (see 

property values listed in Table 1). The aluminum outer skin and frames were assigned properties of 

Mat 24, an elastic-plastic material model. The steel bolts were simulated using beam elements that 

were assigned material properties of hardened steel. 

A depiction of the model of the conusoid energy absorber is shown in Figure 20(a). The conusoid 

was constructed of four layers of hybrid carbon-Kevlar® plain weave fabric with a stacking 

sequence of [+45°/-45°/-45°/+45°]. The individual ply orientations and thicknesses were input 

using the *PART_COMPOSITE command in LS-DYNA®. A nominal element edge length of 0.3-

in. was used in the conusoid mesh. The finite element model of the sinusoid subfloor is shown in 

Figure 20(b). As before, the face sheets were represented as two layers of hybrid carbon-Kevlar®

plain weave fabric with a stacking sequence of [±45°] that were assigned Mat 58 material properties 

(see Table 1). The solid elements representing the foam core were assigned Mat 63. A nominal 

element edge length of 0.25-in. was used in the sinusoid mesh. Note that in the test article, the 

energy absorbers were attached to the outer skin and floor using rivets. In the model, the rivets 

were not physically modeled; however, coincident nodes were used to tie the parts together. 

An automatic single surface contact was assigned to the model with static and dynamic coefficients 

of friction of 0.3. This general contact definition is used to prevent any node from penetrating any 

element within the model. The model was executed using LS-DYNA® SMP version 971 on a 

Linux-based workstation with 8 processors and required 31.75 hours of clock time to execute the 

simulation for 0.065-seconds. Model output included time-history responses of the 320- and 681-lb 

lumped masses, and image sequences of structural deformation. 

 Figure 19. Finite element model of the CH-46E barrel section. 
16 

 (a) Conusoid finite element model. (b) Sinusoid foam sandwich model. 

Figure 20. Depictions of the finite element models of the two energy absorbing subfloors. 

5.3 Barrel Section Test-Analysis Comparisons 

Test-analysis time history responses of the 681-lb mass, located above the conusoid energy 

absorber, are plotted in Figure 21. The predicted acceleration response shows a spike that occurs at 

0.032-s. This spike is attributed to bottoming out of the subfloor section, meaning that the floor 

impacts the outer skin of the fuselage section. The “bottoming out” response was confirmed when 

viewing the high-speed video. The test acceleration response also indicates that the subfloor 

bottomed out, as evident by the gradual increase in acceleration leading to a 34.5-g peak near the 

end of the pulse. Average accelerations of 15.0- and 20.3-g were calculated for the test and 

predicted responses, respectively, for a pulse duration of 0.0- to 0.06-s. The selection of the pulse 

duration for calculating average acceleration was made based on the time at which both acceleration 

traces were approaching zero magnitude. The velocity responses indicate that the model is too stiff 

and predicted velocity is removed much more quickly than for the test. In addition, the model 

predicts a much higher rebound velocity than the test, which indicates that the model contains too 

much elastic energy and that the unloading response is not adequately captured. The predicted 

maximum crush displacement is 7.8-in. and the experimental maximum crush displacement is 8.67-

in., a difference of approximately 0.9-in. Finally, it should be noted that the average test 

acceleration response falls below the design goal of 25- to 40-g. 

 (a) Acceleration responses. (b) Velocity responses. (c) Displacement responses. 

Figure 21. Comparisons of test-analysis responses of the 681-lb mass over the conusoid subfloor. 
17 

A time sequence of predicted deformations is shown in Figure 22 for the conusoid subfloor beam. 

Initially, the crush pattern is non-uniform due to the fact that the 681-lb mass is actually attached to 

the floor using two I-beams separated by 26-in. with flange widths of 6-in. The sides of the 

conusoid subfloor that attach to the fuselage frames are relatively undamaged. Finally, a post-test 

photograph of the conusoid subfloor, which was removed from the test article, is shown in Figure 

23. The deformation and failure pattern is more uniform, especially in the central portion of the 

subfloor than predicted. It is also interesting to note the permanent deformation pattern of the 

polycarbonate floor above the conusoid. 

(a) Time=0.01-s. 

(b) Time=0.02-s.

(c) Time=0.03-s. 

(d) Time=0.04-s. 

Figure 22. Predicted deformation pattern of the conusoid subfloor. 

Figure 23. Photograph of the conusoid energy absorber, removed from the test article. Note that the 

conusoid is lying on its side with a portion of the polycarbonate floor still attached. 

Test-analysis time-history responses of the 320-lb mass, located above the sinusoid energy 

absorber, are plotted in Figure 24. While the predicted responses demonstrate reasonable 

comparison with test, the model fails to predict the large increase in acceleration that occurs just 

prior to 0.03-s. This 64-g peak is attributed to impact of the steel bolts with the outer skin. Even 

though the model includes the bolts, the predicted acceleration response does not match the test. 

Average accelerations of 14.2- and 17.0-g were calculated for the test and predicted responses, 

respectively, for a pulse duration of 0.0- to 0.0575-s. It should be noted that average test and 

predicted accelerations are well below the design goal of 25- to 40-g. The test-analysis velocity 

responses are shown in Figure 24(b), both of which cross zero at the same time (0.0326-s), even 
18 

though the test and predicted curves deviate shortly after 0.01-s. The model predicts a much higher 

rebound velocity than the test, which indicates that the model returns too much elastic energy and 

that the unloading response is not adequately represented. The predicted maximum crush 

displacement is 5.24-in. and the experimental maximum crush displacement is 6.3-in., a difference 

of approximately 1-in. 

 (a) Acceleration responses. (b) Velocity responses. (c) Displacement responses. 

Figure 24. Comparisons of experimental and predicted responses of the 320-lb mass. 

Three images from the high-speed video are compared with predicted model deformation in Figure 

25 at times of 0.01-, 0.03-, and 0.043-s. By 0.043-s, fairly severe deformations are evident in the 

upper portion of the fuselage section. This deformation is asymmetric, located on one side only, 

and is not captured by the model. While the model exhibits some oscillatory motion in the outer 

skin, it does not show the large deformation evident in the test. It should be noted that the 

deformation in the test article is elastic, since the upper fuselage section returns to its normal shape 

by the end of the pulse, as indicated in the post-test photograph shown in Figure 17(a). 

A time sequence of predicted sinusoid foam sandwich deformations is shown in Figure 26. 

Initially, the crush pattern is non-uniform due to the fact that the subfloor is loaded by two discrete 

masses, the 320-lb mass on one side and the seat and occupants on the other. Note that the sides of 

the sinusoid subfloor that attach to the fuselage frames are relatively undamaged. Finally, a posttest photograph of the sinusoid subfloor, which was removed from the test article, is shown in 

Figure 27. The deformation and failure pattern of the sinusoid is more uniform than the predicted 

response; however, the locations and types of damage are similar. 

Time, s Test Model 

0.01-s 
19 

0.03-s 

0.043-s 

Figure 25. Comparison of experimental and predicted deformation patterns. 

(a) Time=0.0-s.

(b) Time=0.01-s 

(c) Time=0.02-s.

(d) Time=0.03-s 

(e) Time=0.04-s 

Figure 26. Predicted deformation pattern of the sinusoid foam sandwich subfloor. 
20 

Figure 27. Photograph of the sinusoid foam sandwich energy absorber, removed from the test 

article. Note that the subfloor section is lying on its side with a portion of the polycarbonate floor 

still attached. 

6.0 FULL-SCALE CRASH TESTING AND SIMULATION OF TRACT 2 

6.1 TRACT 2 Full-Scale Crash Test 

A second CH-46E helicopter airframe was prepared for crash testing and loaded in a similar manner 

as the TRACT 1 test article [2]. In addition, the TRACT 2 aircraft was retrofitted with three 

different composite energy absorbing subfloor concepts [3]. The shear panel at FS220 was replaced 

with a corrugated web energy absorber developed by the German DLR and the Australian ACSCRC and fabricated of graphite fabric material [4, 5]. The shear panel at FS254 was replaced with 

the sinusoid foam sandwich energy absorber and the shear panel at FS286 was replaced with the 

conusoid energy absorber [6]. A photograph showing the retrofit of the sinusoid and conusoid 

energy absorbers into the TRACT 2 subfloor is shown in Figure 28. Unlike the barrel section, the 

original floor in the CH-46E was not replaced with polycarbonate material. However, for viewing 

of the subfloor response during the crash test, rectangular-shaped windows were cut into the floor 

panels at discrete locations between the seat rails and polycarbonate was used to fill these openings. 

Photographs of the conusoid and sinusoid energy absorbers are shown in Figures 29(a) and (b), 

respectively, as viewed through the polycarbonate windows. The energy absorbers were 63-in. 

wide, 9.2-in. tall, and 1.5-in. deep. The conusoid and sinusoid subfloor beams weighed 4.02- and 

4.52-lb., respectively. In comparison, the aluminum shear panel that was replaced weighed 5.75-lb. 

Figure 28. Photograph showing the retrofitted sinusoid and conusoid energy absorbers. 
21 

On October 1, 2014, the TRACT 2 full-scale crash test was conducted at the LandIR Facility. Preand post-test photographs are shown in Figure 30. Nine organizations, including NASA, NAVAIR, 

DLR/ACS-CRC, FAA, US Army Aeromedical Research Laboratory (USAARL), US Army 

CARGO, Cobham Life Support/BAE Systems, and Safe Inc., took part in the TRACT 2 activity, 

contributing 18 experiments related to occupant seating and restraints, composite crashworthiness, 

and emergency locator transponder survivability, as described in References 3 and 24-27. The 

TRACT 2 test article was instrumented with over 360 data channels, including 13 ATDs, 12 onboard high-speed cameras, 10 on-board Go-Pro cameras, and 12 external high-speed cameras. Data 

were recovered from over 95% of the channels, with only one on-board high-speed camera and two 

Go-Pro cameras failing to operate. Measured impact conditions were 403.8-in/s (33.65-ft/s) 

forward velocity and 304.32-in/s (25.36-ft/s) vertical velocity. The airframe attitude at impact was 

2.6° pitch (nose up) and 3.5° roll (left side down). The total weight of the test article was 10,534-lb. 

 (a) Conusoid energy absorber. (b) Sinusoid energy absorber. 

Figure 29. Photographs of the installed energy absorbers. 

 (a) Pre-test photograph. (b) Post-test photograph. 

Figure 30. Pre- and post-test photographs of TRACT 2. 

The impact surface was a sand/clay mixture. Soil characterization testing [28, 29] was performed 

both prior to and following the crash test to determine soil density, moisture content, bearing 

strength, and California Bearing Ratio (CBR). The measured moisture content of the soil was high 

and it ranged from 9.7 to 16%. Based on post-test measurements, the airframe skidded along the 
22 

soil surface approximately 51-in. in the forward direction following initial contact. A post-test 

photograph is shown in Figure 31 showing the slide out distance from the rear of the airframe. 

 

Figure 31. Post-test photograph taken from the rear of the airframe looking forward. 

During the impact, the outer belly skin buckled and tore between FS220 and FS286 as it plowed 

through the soil. The bottom skin skidded approximately 51-in. along the surface of the soil, 

leaving an 8- to 9-in.-deep divot (maximum depth). As the outer belly skin failed, the floor 

continued to move forward, which produced shearing in the subfloor beams. The outer skin was 

torn in several places, while the composite energy absorbing subfloor beams rotated globally under 

shear loading without significant crushing, as shown in the photograph of Figure 32. The severe 

outer skin deformation and failure is attributed to wet soil conditions, measured to have a variable 

moisture content. The crash test was performed a few days following a rainstorm. Even though the 

soil was covered prior to the test, water was able to penetrate a seam in the tarp. The moist soil 

produced a higher than anticipated coefficient of friction. For example, TRACT 1 was tested under 

the same impact conditions onto the same soil and had a slide out of approximately 96-inch [2]. 

Figure 32. Post-test photograph of the outer belly skin deformation between FS220 and FS286. 
23 

Following the impact test, the composite subfloors were removed from the test article. Photographs 

of the conusoid and the sinusoid foam sandwich energy absorbers are shown in Figures 33(a) and 

(b), respectively. Note that both photographs show the energy absorbers as they would be 

positioned facing rearward. The rearward side of both energy absorbers was painted and marked 

for camera viewing. The conusoid exhibited fracturing on the left and right sides where the energy 

absorber attached to the fuselage frames. No evidence of crushing or plastic deformation was 

observed. The sinusoid displayed areas of crush initiation, especially on the bottom left side, as 

highlighted in the inset photograph of Figure 33(b). Note that a 600-lb. mass was attached to the 

floor on the left side that straddled the sinusoid energy absorber at FS254. However, the amount of 

crushing was estimated to be less than 0.5-in. 

(a) Post-test photograph of the conusoid energy absorber, rear view looking forward. 

 (b) Sinusoid energy absorber, rear view looking forward. Inset shows regions of crush initiation. 

Figure 33. Post-test photographs of two composite energy absorbers. 

6.2 TRACT 2 Finite Element Model 

Development of a finite element model of the TRACT 2 test article was completed and predictions 

of structural impact responses were generated. The airframe model is shown in Figure 34. The 

model consists of: 218,251 nodes; 13,178 beam elements; 102,413 Belytschko-Tsay shell elements; 

119,632 solid elements; 473 parts; 27 material properties; 34 element masses; 19 constrained nodal 

rigid bodies; 1 initial velocity card; and 1 body load representing gravity. The composite shell 

elements forming the conusoid and the face sheets of the sinusoid foam sandwich energy absorbers 
24 

were represented using Mat 58, with properties listed in Table 1. The foam in the sinusoid 

sandwich was represented using solid elements assigned Mat 63. Due to early problems with 

negative volumes in the solid elements representing the sinusoid foam, an erosion card was added to 

eliminate solid elements once they reached a volumetric strain of 0.9. 

Figure 34. Depiction of the TRACT 2 finite element model. 

Finite element models of the sinusoid and conusoid energy absorbers were incorporated into the 

TRACT 2 model, as shown in Figure 35. These subfloors were located at FS254 and FS286, 

respectively. Nominal shell element edge length for the conusoid was 0.3-in., compared with a 0.25-

in. element edge length for the sinusoid. A detailed structural model of the DLR/ACS-CRC 

corrugated web subfloor, located at FS220, was not included and was instead modeled as a typical 

aluminum shear panel. 

The soil was represented using solid elements that were assigned Mat 5 

(*MAT_SOIL_AND_FOAM) in LS-DYNA®, which is a material model for representing soil and 

foam [30]. The soil block was 24-in. deep x 148-in. wide x 600-in. long. A coefficient of friction 

of 0.5 between the airframe and the soil was used in an automatic single surface contact definition. 

Initially, the soil was represented as a single block with one material model assigned; however, 

based on the soil characterization results, the model was changed to a layered soil configuration. 

The top 3-in.-deep layer of soil was represented using Mat 5 with input properties obtained from 

soil tests conducted on gantry unwashed soil, which were performed for the Orion program [28]. 

The bottom 21-in. deep layer was also represented using Mat 5 with input properties of sand, whose 

bearing strength matched in-situ test results conducted prior to and after the TRACT 2 crash test. 

The bottom and side nodes of the soil model were constrained from motion using a SPC definition 

in LS-DYNA®. 

All nodes forming the helicopter airframe were assigned measured initial conditions of 403.8-in/s 

(33.65-ft/s) forward and 304.32-in/s (25.36-ft/s) vertical velocities. In addition, the TRACT 2 

model was oriented to match the measured impact attitude (2.6° nose-up pitch, 3.5° left-side down 

roll). Seat/occupant and discrete masses, such as the cargo experiment in the rear of the airframe, 

were represented using Constrained Nodal Rigid Bodies (CNRBs), as illustrated in Figure 36(a). A 

600-lb lumped mass was attached to the left-side floor, centered about FS254, and located above the 

sinusoid energy absorber, as shown in Figure 36(b). 
25 

Figure 35. Picture of the outer shell and subfloor sections in the final model. 

(a) CNRB locations using in the model. 

(b) Location of 600-lb concentrated mass over the sinusoid energy absorber. 

Figure 36. Locations of CNRBs and concentrated masses used in the model. 
26 

Prior to executing the model, a separate simulation was performed, without the soil, to obtain 

weight and balance information on the airframe model. Experimental and predicted data are listed 

in Table 2. In general, the inertial properties of the model compare well with test data. The 

simulation was executed using LS-DYNA® Version 971 R6.1.1 SMP (double precision) for 0.1-s, 

which required 74 hours and 35 minutes of CPU time on a Linux-based workstation computer with 

8 processors. 

Table 2. Comparison of model and test weight and balance data. 

Test Model 

Weight, lb. 10,534 10,534 

CGx, in. 262.8 269.6 

CGy, in. ±0.5 -0.91 

CGz, in. -10.0 -9.56 

Nodal output was requested at locations corresponding to accelerometers mounted on the cabin 

floor in the test article. The locations of floor-mounted accelerometers, which were attached at the 

frame/floor junctions on both the left and right sides of the airframe, are illustrated in Figure 37. 

Test-analysis comparisons were generated at these locations and the experimental and predicted 

responses were filtered using a 4-pole Butterworth low-pass filter with a cutoff frequency of 60-Hz.

Figure 37. Schematic of fuselage section showing floor-mounted accelerometer locations. 

6.3 TRACT 2 Test-Analysis Comparisons 

Test and analysis acceleration time histories are shown in Figures 38 through 45 for FS160 through 

FS410. As noted in Figure 37, these accelerometers were located at the intersection of the floor and 

the fuselage frames, on the far edges of the floor. In general, the test-analysis comparison results 

demonstrate a reasonable level of agreement, especially when considering the fact that the actual 

deformation and failure behavior of the outer skin between FS220 and FS286 was not captured by 

the model. In the test, the outer belly skin buckled and tore as it plowed through the soil. It is 

speculated that high shear loading caused the energy absorbing subfloor beams to rotate globally 

without crushing, as shown in Figure 32, based on the loss of their upper and lower attachments to 

the floor and outer skin. In order to replicate this response, the energy absorbing subfloor beams in 

the model would have to be separated from the outer skin and the floor, allowing them the freedom 
27 

to rotate in a global manner as their top and bottom constraints failed. However, these 

modifications to the model were not attempted. 

 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 38. Forward and vertical test and analysis acceleration time histories at Left Side FS160. 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 39. Forward and vertical test and analysis acceleration time histories at Right Side FS160. 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 40. Forward and vertical test and analysis acceleration time histories at Right Side FS220. 
28 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 41. Forward and vertical test and analysis acceleration time histories at Left Side FS254. 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 42. Forward and vertical test and analysis acceleration time histories at Right Side FS286. 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 43. Forward and vertical test and analysis acceleration time histories at Right Side FS350. 
29 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 44. Forward and vertical test and analysis acceleration time histories at Left Side FS410. 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 45. Forward and vertical test and analysis acceleration time histories at Right Side FS410. 

In the model, the composite subfloor beams behaved in an ideal fashion and exhibited stable 

crushing. The outer belly skin stayed intact and the energy absorbers crushed between the floor and 

the outer skin. The conusoid subfloor crushed 48.8% of its original 9.2-in. height, with maximum 

crushing occurring at 0.06-s. The sinusoid subfloor crushed 42.6% of its original height with 

maximum crush displacement occurring at 0.06-s. Fringe plots of vertical displacement are shown 

for the conusoid and sinusoid energy absorbers in Figures 46(a) and (b), respectively, at the time of 

maximum vertical crush displacement. Of course, these predictions were not realized during the 

test.

Test-analysis comparison results are shown in Figures 47(a) and (b) for two accelerometers that 

recorded forward and vertical responses at the base of the double seat located over the conusoid at 

FS286, near the center of the floor. The experimental and analytical forward acceleration responses 

are similar and they match the results previously documented in Figures 38(a) through 45(a) for the 

floor/frame intersection regions. However, the vertical acceleration results vary considerably from 

previous results. The experimental trace exhibits a dramatic initial peak of 60-g, which is higher 
30 

than previously shown vertical acceleration traces, followed by a drop in acceleration and a 

subsequent peak of 22-g. The response is indicative of a sudden shock experienced by the 

accelerometer, as might be caused by fracture of the sides of the conusoid energy absorber from the 

fuselage frame at FS286. In contrast, the predicted response exhibits a stable, fairly uniform 

acceleration of low magnitude (approximately 20-g). This response is indicative of stable crushing 

of the conusoid, which was predicted, as shown in Figure 46(a). 

(a) Conusoid energy absorber. 

(b) Sinusoid energy absorber. 

Figure 46. Depictions of the maximum vertical crush displacements of the conusoid and sinusoid 

energy absorbers in TRACT 2 at t=0.06-s. 

 (a) Forward acceleration. (b) Vertical Acceleration. 

Figure 47. Forward and vertical test and analysis acceleration time histories at FS286. 
31 

A fringe plot of z-displacement is shown in Figure 48(a) of the soil deformation at the time of 

maximum predicted displacement, which was 5.325-in. Measured maximum displacements were 

recorded between 8- and 9-in. A photograph of the post-test soil crater is shown in the photograph 

of Figure 48(b). Note that due to the left roll angle, the soil crater is asymmetric about the 

centerline. The two images are similar and highlight the asymmetry of the impact event. 

 (a) Fringe plot of z-displacement. (b) Overhead photograph of soil crater. 

Figure 48. Fringe plot of predicted soil deformation and post-test photograph. 

7.0 DISCUSSION OF RESULTS 

Two objectives of the TRACT 2 project were to assess the capabilities of two novel composite 

energy absorbers under dynamic impact loading, and to investigate the capabilities of LS-DYNA®

to predict the experimental behavior. As stated previously, the design goals for the energy 

absorbers were to achieve between 25- to 40-g average crush accelerations, to minimize peak crush 

loads, and to generate relatively long crush stroke limits under dynamic loading conditions, typical 

of those experienced during the TRACT 1 [2] full-scale crash test. The results of the component 

tests proved that the energy absorbers met the goal with average crush accelerations of 28-g for the 

conusoid and 21.8-g for the sinusoid. The lower average crush acceleration for the sinusoid 

component simply translates into a larger crush stroke than was seen for the conusoid. Testanalysis comparisons are listed in Table 3 for each level of evaluation, component through full-scale 

testing. Results are presented as percentage differences of average crush acceleration and 

maximum crush displacements, as appropriate. Note that component test-analysis comparisons 

agree within 13%. 

As shown in Table 3, barrel section comparisons are not as good as for the component tests, mainly 

due to the fact that some anomalies occurred during the drop test. The conusoid energy absorber 

bottomed out and two steel bolts interfered with crushing of the sinusoid. However, despite these 

anomalies, test-analysis comparisons were within 20%, with one exception. For the TRACT 2 test, 

only average acceleration comparisons are listed in Table 3 for forward (Fwd) and vertical (Vt) 

acceleration responses generated from floor-mounted accelerometers located between FS160 and 

FS410. For these comparisons, the pulse durations used to determine average accelerations are 

listed in parentheses. The selection of the pulse duration for calculating average acceleration was 

made based on the time at which both test and analysis acceleration traces were approaching zero 

magnitude. The test-analysis comparisons range from excellent (2.1% agreement for the right-side 
32 

vertical acceleration response at FS410) to poor (81.3% agreement for the left-side vertical 

acceleration response at FS160). For the vertical acceleration comparisons, the level of agreement 

for six of eight locations was within 18%. It is also useful to note that the measured average 

forward accelerations were consistent, ranging between 5.8- to 7.8-g, if the one data point obtained 

from the left-side floor at FS160 is ignored. Finally, it is important to note that, based on soil 

anomalies during the crash test, a true assessment of conusoid and sinusoid energy absorption 

behavior as retrofit concepts could not be performed. 

Table 3. Test-analysis comparisons. 

Test Article Description Metric Test Model % Difference 

Conusoid Component Avg. Acceleration, g 

(0.03-s pulse duration) 

28.0 28.2 1.43 

Max crush displacement, in. 2.9 2.53 12.8 

Sinusoid Component Avg. Acceleration, g 

(0.03-s pulse duration) 

21.8 22.9 5.0 

Max crush displacement, in. 4.0 3.8 5.0 

Barrel Section, 681-lb mass, 

located above the conusoid 

Avg. Acceleration, g 

(0.06-s pulse) 

15 17.5 16.7 

Max crush displacement, in. 8.67 7.8 10.0 

Barrel section, 320-lb mass, 

located above the sinusoid 

Avg. Acceleration, g 

(0.06-s pulse) 

14.2 17.0 19.7 

Max crush displacement, in. 6.3 5.24 16.8 

TRACT 2, FS160, Lt Fwd Avg. Accel., g (0.15-s) 3.75 6.8 81.3 

TRACT 2, FS160, Lt Vt Avg. Accel., g (0.078-s) 8.5 11.9 40.0 

TRACT 2, FS160, Rt Fwd Avg. Accel., g (0.133-s) 6.3 4.4 30.2 

TRACT 2, FS160, Rt Vt Avg. Accel., g (0.1-s) 8.7 8.3 4.6 

TRACT 2, FS220, Rt Fwd Avg. Accel., g (0.13-s) 5.9 4.4 25.4 

TRACT 2, FS220, Rt Vt Avg. Accel., g (0.08-s) 13.0 11.7 9.9 

TRACT 2, FS254, Lt Fwd Avg. Accel., g (0.14-s) 6.0 4.1 31.7 

TRACT 2, FS254, Lt Vt Avg. Accel., g (0.1-s) 10.8 10.3 4.3 

TRACT 2, FS286, Rt Fwd Avg. Accel., g (0.15-s) 6.4 4.0 37.5 

TRACT 2, FS286, Rt Vt Avg. Accel., g (0.09-s) 9.0 9.9 10.0 

TRACT 2, FS286, Center Fwd Avg. Accel., g (0.15-s) 5.7 3.9 31.6 

TRACT 2, FS286, Center Vt Avg. Accel., g (0.09-s) 9.7 11.9 22.7 

TRACT 2, FS350, Rt Fwd Avg. Accel., g (0.15-s) 7.1 3.9 45.1 

TRACT 2, FS350, Rt Vt Avg. Accel., g (0.08-s) 8.0 11.6 45.0 

TRACT 2, FS410, Lt Fwd Avg. Accel., (0.15-s) 5.8 3.8 34.5 

TRACT 2, FS410, Lt Vt Avg. Accel., g (0.09-s) 10.7 12.6 17.8 

TRACT 2, FS410, Rt Fwd Avg. Accel., g (0.09-s) 7.8 6.3 19.2 

TRACT 2, FS410, Rt Vt Avg. Accel., g (.07-s) 14.0 14.3 2.1 

8.0 CONCLUSIONS 

This paper has described the development and multi-level evaluation of two composite energyabsorbing concepts. Both concepts were designed to achieve between 25- to 40-g sustained average 

crush accelerations, to minimize peak crush loads, and to generate relatively long crush stroke limits 
33 

during retrofit onto the second Transport Rotorcraft Airframe Crash Testbed (TRACT 2) full-scale 

crash test. The first concept, designated the “conusoid,” is a conusoidal-shaped design based on 

alternating right-side-up and up-side down half-cones placed in a repeating pattern. The conusoid 

combines a simple cone design, with sinusoidal beam geometry to create a structure that utilizes the 

advantages of both configurations. The conusoid was fabricated of four layers of hybrid graphiteKevlar® fabric with layers oriented at [+45°/-45°/-45°/+45°]. The second energy absorber, 

designated the “sinusoid,” is a sinusoidal foam sandwich design, which consists of two face sheets 

oriented at [±45°] fabricated of hybrid graphite-Kevlar® fabric material with a 1.5-in. closed-cell 

foam core separating the face sheets. 

The energy absorbing concepts were evaluated using a multi-level, building-block approach, 

including both testing and LS-DYNA® simulation. Initially, component specimens were subjected 

to vertical impact using a 14-ft. drop tower. The components had nominal dimensions of 12-in. in 

length, 7.5-in. in height and an overall width of 1.5-in. The component tests were used to assess the 

energy absorption capabilities of the two composite designs. The impact condition for all of the 

dynamically crushed specimens was approximately 264-in/s (22-ft/s). 

Next, subfloor beams of the conusoid and sinusoid foam sandwich configurations were 

manufactured and retrofitted into a barrel section of a CH-46E helicopter airframe. A vertical drop 

test of the barrel section was conducted at 297.6-in/s (24.8-ft/s) onto concrete. The objectives of the 

test were to evaluate: (1) the performance of the two energy absorbers during a full-scale drop test 

prior to the TRACT 2 test, (2) the fabrication methods for the energy absorbers, (3) the structural 

integrity of the retrofit, (4) the strength of the polycarbonate floor, and (5) imaging techniques used 

during the test. 

Finally, the two energy absorbers were retrofitted into the subfloor of the TRACT 2 test article, a 

full-scale CH-46E helicopter airframe. A full-scale crash test was performed onto soil with impact 

conditions of 403.8-in/s (33.65-ft/s) forward and 304.32-in/s (25.36-ft/s) vertical velocity. The test 

article contained numerous onboard experiments; however, a major goal of the test was to evaluate 

the performance of novel composite energy absorbing subfloor designs for improved 

crashworthiness. 

For each multi-level evaluation of the subfloor concepts, test data were used to validate simulations 

performed using the commercial nonlinear explicit, transient dynamic finite element code, LSDYNA®. Finite element models were developed to represent dynamic crushing of the component 

specimens, the vertical impact of the retrofitted barrel section, and the retrofitted TRACT 2 airframe 

impacting onto soil. 

Major findings of this research effort are listed, as follows: 

• Both the conusoid and sinusoid foam sandwich concepts proved to be excellent energy 

absorbers, as demonstrated through component impact tests. The conusoid component 

exhibited stable folding and crushing up to 38.7% stroke with an average acceleration of 

28.0-g, thus meeting the stated design goal. Likewise, the sinusoid component absorbed 

energy through localized uniform folding of the face sheets and foam crushing. An average 

acceleration of 21.8-g was recorded for the sinusoid over 50% crush stroke. 

• For both components, the LS-DYNA® predictions showed excellent comparison with test 

data. The LS-DYNA® model of the conusoid predicted an average acceleration of 28.4-g for 
34 

the conusoid. Likewise, the average acceleration of the predicted response for the sinusoid 

component is 22.9-g. 

• The barrel section drop test results were complicated by the fact that the conusoid energy 

absorber bottomed out, allowing the floor to impact the outer skin. In addition, two long 

bolts used to attach the concentrated mass to the floor over the sinusoid were untrimmed, 

allowing the bolts to impact the outer skin and deform plastically. Both of these events 

resulted in large increases in the acceleration responses near the end of the pulses, as 

measured on the two concentrated masses located over the conusoid and sinusoid energy 

absorbers. Despite these complications, average accelerations of 15.0- and 14.2-g were 

measured on the 681-lb and 320-lb concentrated masses located over the conusoid and the 

sinusoid energy absorbers, respectively. 

• During the barrel section impact, the conusoid energy absorber exhibited 58% crush stroke 

and displayed fracturing and delamination of the hybrid carbon-Kevlar® walls. The 

sinusoid foam sandwich energy absorber exhibited 49.3% crush stroke and displayed 

crushing of the foam core, and fracturing of the face sheets starting from the bottom, curved 

edge. The sinusoid energy absorber in the barrel section did not exhibit the uniform folding 

pattern seen in the face sheets of the component specimen. 

• LS-DYNA® model predictions for the barrel section drop test were reasonable; however, 

results indicated that the model was generally too stiff. For example, the predicted 

maximum crush displacement of the conusoid energy absorber was 7.8-in. and the 

experimental maximum crush displacement was 8.67-in., a difference of approximately 0.9-

in. Likewise, the predicted maximum crush displacement of the sinusoid was 5.24-in. and 

the experimental maximum crush displacement was 6.3-in., a difference of 1-in. 

• Results from the TRACT 2 full-scale crash test were also complicated by anomalies. Moist 

soil increased the coefficient of friction and reduced the stopping distance of the test article 

by half, compared with the TRACT 1 test. Due to excessive damage of the outer belly skin, 

the composite energy absorbers failed to crush and rotated globally as they became 

separated from the floor and outer skin. Regardless, over 95% of 350-channels of data were 

collected during the impact test. 

• Test-analysis comparisons with data obtained from accelerometers, mainly located at the 

floor/fuselage frame intersections at FS160 to FS410, were evaluated. These results 

demonstrated reasonable test-analysis comparisons, despite the fact that the model predicted 

idealized crush responses for both energy absorbers, which did not match the observed 

behavior. 

• Finally, based on soil anomalies during the crash test, a true assessment of conusoid and 

sinusoid energy absorption behavior as retrofit concepts could not be performed. 

 

9.0 ACKNOWLEDGEMENTS 

The authors of this report gratefully acknowledge personnel from the Navy Flight Readiness Center 

in Cherry Point, NC for providing the TRACT 2 CH-46E airframe. In addition, Navy personnel 

provided a baseline NASTRAN model that was used as the baseline to generate the LS-DYNA® 

finite element model. We would also like to thank engineers, technicians and contractor personnel 

who work at the NASA Langley LandIR Facility for instrumenting and configuring the test article, 

collecting test data, taking photographs and videos, and performing soil characterization testing. 
35 

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May 27-29, 2009. 
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Multi-Level Experimental and Analytical Evaluation of Two Composite 

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14. ABSTRACT

Through multi-level testing and simulation performed under the Transport Rotorcraft Airframe Crash Testbed (TRACT) research program. A conical-shaped 

energy absorber, designated the conusoid, was evaluated that consisted of four layers of hybrid carbon-Kevlar® plain weave fabric oriented at [+45°/-45°/-45°

/+45°] with respect to the vertical, or crush, direction. A sinusoidalshaped energy absorber, designated the sinusoid, was developed that consisted of hybrid 

carbon-Kevlar® plain weave fabric face sheets, two layers for each face sheet oriented at ±45° with respect to the vertical direction and a closed-cell ELFOAM

® P200 polyisocyanurate (2.0-lb/ft3) foam core. The design goal for the energy absorbers was to achieve average floor-level accelerations of between 25- and 

40-g during the full-scale crash test of a retrofitted CH-46E helicopter airframe, designated TRACT 2. Variations in both designs were assessed through 

dynamic crush testing of component specimens. Once the designs were finalized, subfloor beams of each configuration were fabricated and retrofitted into a 

barrel section of a CH-46E helicopter. A vertical drop test of the barrel section was conducted onto concrete to evaluate the performance of the energy 

absorbers prior to retrofit into TRACT 2. The retrofitted airframe was crash tested under combined forward and vertical velocity conditions onto soil, which is 

characterized as a sand/clay mixture. Finite element models were developed of all test articles and simulations were performed using LSDYNA ®, a 

commercial nonlinear explicit transient dynamic finite element code. 

15. SUBJECT TERMS

Composite airframe structures; Composite continuum damage; Composite progressive failure model; Element modeling; 

Explicit transient dynamic finite; Impact testing; LS-DYNA; Mechanics model; Nonlinear

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