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Investigation of a crash concept for CFRP transportaircraft based on tension absorptionPaul Schatrowa & Matthias Waimera

a German Aerospace Center (DLR), Institute of Structures and Design, Stuttgart, GermanyPublished online: 19 May 2014.

To cite this article: Paul Schatrow & Matthias Waimer (2014) Investigation of a crash concept for CFRP transport aircraftbased on tension absorption, International Journal of Crashworthiness, 19:5, 524-539, DOI: 10.1080/13588265.2014.917498

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Investigation of a crash concept for CFRP transport aircraft based on tension absorption

Paul Schatrow* and Matthias Waimer

German Aerospace Center (DLR), Institute of Structures and Design, Stuttgart, Germany

(Received 7 October 2013; accepted 21 April 2014)

Transport aircraft made of carbon fibre reinforced plastics (CFRP) have to provide an equivalent crashworthiness comparedto today’s aluminium aircraft designs. However, CFRP structures typically show brittle failure behaviour under complexloading conditions and little energy absorption, whereas aluminium structures provide comparably high energy absorptiondue to their ductile failure characteristics. Improved crashworthiness for CFRP fuselages can be obtained by theinstallation of special crash devices, which are designed for energy absorption by progressive failure in compression,tension or bending. The realisation of crashworthy CFRP fuselage designs with the focus on compression or bendingabsorbers is often associated with significant mass penalty compared to the purely static sizing of the correspondingfuselage structure. In this context, an alternative crash kinematics was developed and numerically investigated in whichmost of the kinetic energy is dissipated by tension absorption in the sub-cargo area of a fuselage structure. The numericalstudy was performed on the basis of a purely vertical impact with a two-bay fuselage section using the explicit finiteelement (FE) solver Abaqus/Explicit. The simulation results show for the developed crash kinematics several advantages,e.g. reduced mass penalty, with the tension absorption concept compared to crash concepts that use energy absorption byprogressive crushing in the sub-cargo area.

Keywords: CFRP transport aircraft; crashworthiness; tension absorption; crash simulation; Abaqus/Explicit; kinematicsmodel

Introduction

In today’s commercial transport aircraft aluminium is the

most used material due to its favourable strength to weight

ratio [19]. The lightweight design and consequently the effi-

ciency of future transport aircraft can be further improved

by replacing aluminium alloys with composite materials,

respectively, composite designs, particularly with respect to

aircraft primary structures (B787, A350XWB).

The main challenge of using carbon fibre reinforced

plastics (CFRP) for crashworthy aircraft structures is the

brittle failure behaviour compared to aluminium alloys

and its propensity towards uncontrolled failure with little

energy absorption under complex loading conditions,

which typically occur in an aircraft crash [6]. Safety regu-

lations require for CFRP transport aircraft an equivalent

crashworthiness compared to nowadays transport aircraft,

which are made of aluminium alloys [18]. Sufficient

energy absorption in CFRP transport aircraft can be

achieved by the installation of crash devices in specific

areas of the fuselage structure where substantial failure is

expected due to loading conditions mainly in compression

or bending, and partly in tension.

One possible and well-known absorption concept for

CFRP transport fuselage structures is the progressive

crushing of vertical orientated structural elements which

are located in the sub-cargo area. Potential design solu-

tions of this concept are given in [4,8,12,22,35,42]. How-

ever, the crushing concept in the sub-cargo area requires a

massive backing structure that provides sufficient strength

to sustain the crush forces. This requirement typically leads

to a massive cargo crossbeam and partly frame design with

significantly higher structural mass compared to the static

sizing that does not consider the crash load case.

Besides the sub-cargo area, further kinetic crash

energy can be absorbed in the frame structure. With

respect to typical crash loads, the frame is one of the high-

est loaded primary structures and it significantly contrib-

utes to the total energy absorption, particularly with

respect to the crash displacement that is associated to the

bending failure of the frame structure. Comparably high

bending rotations on high moment levels are necessary to

achieve smooth energy absorption during the whole crash

event. Metallic frames made of aluminium or titanium

alloys provide this required failure behaviour solely by

their ductility. Substantially more challenging is this

bending absorption requirement for CFRP frame struc-

tures. In contrast to the progressive crushing mode, CFRP

typically does not provide high mass-specific energy

*Corresponding author. Email: paul.schatrow@dlr.deThis article was originally published with errors. This version has been corrected. Please see Erratum (http://dx.doi.org/10.1080/13588265.2014.930569)

� 2014 Taylor & Francis

International Journal of Crashworthiness, 2014

Vol. 19, No. 5, 524�539, http://dx.doi.org/10.1080/13588265.2014.917498

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absorption in bending failure modes. Brittle rupture is

often the bending failure mode of CFRP frames with lim-

ited energy absorption that is mainly associated with the

elastic energy stored in the frame up to failure. Several

research studies were performed to investigate and to

improve the bending failure mechanisms of CFRP frame

structures. In [5] and [6] single z-shaped CFRP frames as

well as skeleton floor sections with and without skin were

tested statically and dynamically. One outcome of these

tests is the brittle failure behaviour of CRRP frames with-

out significant energy absorption. Improved crashworthy

response of frame structures was investigated in [45] by

optimisation of the cross-sectional dimensions and the

laminate stacking sequence of the frame. The maximum

failure load and partly the amount of energy absorbed by

the frame structure could be improved in this context. To

increase the energy absorption of CFRP frames after first

failure and during high bending rotation angles reference

[41] discusses a research study on hybrid CFRP/titanium

laminates that are used in generic frames to achieve a

more ductile bending failure behaviour. In none of these

research activities on crashworthy CFRP frames sufficient

improvement in the energy absorption under bending

loads could be obtained. It is still a challenge to achieve

equivalent energy absorption in CFRP frames compared

to their metallic equivalents.

In this paper alternative crash kinematics are pre-

sented and discussed whose main absorption concept is

based on tension absorption. The concept is to avoid

energy absorption mechanisms that require massive back-

ing structure (e.g. progressive crushing in the sub-cargo

structure) as well as absorption modes that can hardly be

provided by CFRP structures (e.g. bending absorption in

the frame). In general, tension load benefits a lightweight

structural design and therefore is a promising absorption

mechanism. High-strength backing structure is typically

not required and the absence of instabilities directly leads

to simple designs. An overview of some tension absorber

concepts, e.g. tube expansion, tube inversion and elonga-

tion of a metal tube, is given in [15]. In [34,36,37] an

absorption concept is described where a pin is pulled

through a composite laminate to initiate bearing failure.

This kind of failure mechanism allows the usage of avail-

able aircraft structures, e.g. bolted connections of primary

structures, for tension absorption.

The aim of the preliminary design study presented in

this paper is to investigate potential tension crash concepts

on fuselage section level and to derive potential crash

kinematics for further assessment of these concepts.

Modelling approach for investigation of the tension

crash concept

The presented investigation of a novel crash concept for

CFRP transport aircraft first concentrated on potential

crash kinematics to assess the general crash performance

and to derive failure, respectively, absorption characteris-

tics of different crash absorbers which are installed in the

aircraft fuselage. Besides these crash devices and their

failure characteristics, which are requirements to obtain

the desired crash kinematics, the structural crash loads

and the passenger loads are assessed.

Different FE modelling approaches can be used to

develop and assess new crash concepts on preliminary

design level. The kinematics model approach was applied

in this study, which is described and discussed in the con-

text of other modelling approaches in [39�41]. The fea-

ture of this kinematics modelling is to combine the

benefits of hybrid codes (e.g. DRI-KRASH [33]) and of

detailed FEM techniques. Regions in the fuselage struc-

ture where damage and failure is expected are represented

by macro elements. Other regions which are expected to

remain undamaged are discretised with coarse mesh den-

sity and linear-elastic material formulations. On the one

hand, time expensive calculation processes, such as simu-

lation of crushing or frame bending failure, are repre-

sented by macro elements whose failure characteristics

can be described by force-displacement or moment-rota-

tion curves. The input characteristics are based on realistic

structural failure behaviour and can be obtained from test

results or from full FEM simulations. On the other hand,

in the region of coarse discretisation and linear-elastic

material formulations, the kinematics model approach still

provides sufficient accuracy to represent detailed struc-

tural effects such as frame-skin interaction or instabilities.

Local strains, e.g. in the frame flanges, can be checked

against simplified failure criteria, such as empirical crip-

pling failure data. The main advantage of crash investiga-

tions using the kinematics model approach is the

assessment of absorber input characteristics (force-dis-

placement or moment-rotation) in a crash scenario on

fuselage section level. The influence of different absorber

characteristics on the crash performance of a fuselage sec-

tion can be analysed efficiently. Structural loads and mass

penalty due to crashworthy sizing, loads on the passengers

or the robustness of crash concepts can be investigated

and assessed quickly.

In Figure 1 the fuselage section of a generic full CFRP

transport aircraft is shown, so-called ‘installation areas’

for crash devices are depicted here. On the left side of

Figure 1 the fuselage section is illustrated with potential

crash device concepts for the sub-cargo area, the frame

and the cabin floor. On the right side, corresponding

macro architectures are depicted, which are used in the

kinematics model approach to represent the failure behav-

iour of the crash devices.

The first installation area for the crash devices is

located in the sub-cargo structure. According to the ten-

sion crash concept considered here, kinetic energy shall

be absorbed in this installation area by tensile failure

International Journal of Crashworthiness 525

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processes. Research on aircraft crashworthiness per-

formed in the past indicated comparably high-tension

forces in the cargo crossbeam due to the bending loads

acting on the sub-cargo structure in the first phase of a

crash event [31,32]. Current research activities in cooper-

ation with Airbus Group Innovations consider an innova-

tive cargo floor structure with a filigree crash design for

tension absorption [3]. The physical design of tensile

crash devices in this area is exemplarily depicted by elon-

gation of a metallic structure (energy dissipation by plas-

tic deformation). The physical crash devices are

represented in the simulation model by macro elements

with a force-displacement input characteristic that include

unloading/reloading options. The macro architecture com-

prises two serial connector elements to allow the model-

ling of articulated connection on both ends of the tension

absorber. The connector element is a two-node macro ele-

ment with three translational degrees of freedom at node 1

and three rotational degrees of freedom at node 2.

The second installation area for crash devices is

located in the frame at discrete positions where bending

failure is expected. The macro architecture for modelling

of frame bending failure is called ‘kinematic hinge’. The

positions of the kinematic hinges are defined according to

typical failure locations known from several drop tests

performed in the past [1,2,5�7,9�11,17,20,28�30,44]. A

more accurate positioning of the kinematic hinges,

depending on the specific fuselage design, can be obtained

by evaluating the strain distribution along the frame and

placing the kinematic hinges at positions of local strain

extremum. The physical design of absorber devices in the

frame could be realised e.g. by hybrid CFRP/titanium

laminates as previously described and illustrated in

Figure 1. The macro architecture for kinematic hinges is

described by a cut in the frame whose cross sections are

reinforced by rigid bodies. The fuselage skin remains con-

tinuous. A macro element (connector) describing the

required moment-rotation input characteristic connects

both frame cross sections and is placed in the skin plane

similar to the pivot of fuselage shell post-failure bending

[41]. The connector behaviour can be described with a

main-curve for frame bending absorption and arbitrary

unloading/reloading options.

The third installation area for crash devices is located

in the connection of the frame to the passenger crossbeam

(cabin floor). Here, kinetic energy shall be absorbed by

tensile loads, which act in the passenger crossbeam due to

the so-called ‘ovalisation effect’. The ovalisation effect is

the tendency of the fuselage section, respectively, the

frame structure, to deform to an oval shape due to the

crash loads, which is partly hindered by the stiffness of

the cabin floor. The physical design of tensile crash

Figure 1. Tension crash concept. Left: potential absorption mechanisms. Right: macro architectures.

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devices in connection with the frame to the passenger

crossbeam is exemplarily depicted by bearing failure of a

bolt that is pulled through a laminate. In the kinematics

model the corresponding macro architecture describes

two rigid bodies, which reinforce the end of the passenger

beam as well as the corresponding frame segment. Both

rigid bodies are connected by a macro element that pro-

vides force-displacement and moment-rotation character-

istics with a main-curve and unloading/reloading options.

Besides the local macro models for failure representa-

tion, the kinematics model approach provides an adequate

discretisation for structural regions where failure is not

expected. Due to comparably small deformations in the

scope of the elastic material limits, a coarse mesh size can

be used for discretisation with elements of first-order

shape functions. In addition, structural details that have

minor influence on the overall crash performance (cleats,

mouseholes and rivets) are not modelled in this prelimi-

nary design approach.

The tension crash concept discussed in this article

considers the above described absorption mechanisms

and is focused on maximum energy absorption by tensile

failure mechanisms. The considered crash displacement

is limited by the vertical support struts located between

the frame and the passenger crossbeam. The kinetic

energy of typical crash load cases shall be absorbed up

to the ground contact of these struts. Further absorption

mechanisms in the vertical support struts are not consid-

ered in this study. However, several known concepts

could be used to absorb further energy in the vertical

support struts in crash cases of higher kinetic energy

[14,21].

Detailed investigation of the frame failure modelling

The simulation study on the tension crash concept was

performed on the basis of a generic fuselage design with

CFRP frames of a closed omega-shaped cross section.

Frame failure modelling, using kinematic hinges, is

described in this paragraph with respect to the specific

frame design.

Calibration of the kinematic hinges

The calibration of the kinematics hinges, or more specifi-

cally of the macro element input characteristics, has to be

performed with respect to the stiffness values to obtain

similar elastic behaviour of the kinematic hinges com-

pared to the equivalent and adjacent frame structure. The

stiffness calibration shall consider frame�skin interaction

effects (e.g. effective width of the skin) that highly influ-

ence the bending stiffness and is therefore conducted on

the basis of detailed FE models that represent the corre-

sponding structural parts on a generic level. In Figure 2

different discretisation and modelling levels of the so-

called ‘frame�stringer�skin’ model are shown [40],

which were used for the calibration.

The detailed modelling level represents the frame,

stringers and skin with shell elements using fine discreti-

sation. An element size of 12 mm represents three ele-

ments along the width of the frame flange, which is

sufficiently accurate for representation of the elastic bend-

ing behaviour. Structural details, e.g. mouseholes, are

considered.

Starting with the detailed modelling level (L#1) the

discretisation was simplified stepwise to analyse the influ-

ence of different modelling options that finally lead to the

kinematics modelling level (L#4). The kinematics model-

ling level has to provide the same elastic behaviour as the

detailed modelling level to ensure sufficient accuracy for

frame representation. The basis for comparison of elastic

stiffness is the longitudinal strain along the frame flanges

under pure bending load with positive and negative

moment. The longitudinal strains are determined using

bar elements of negligible stiffness positioned at the frame

flanges.

In Figure 3(a) the strains along the inner frame flange

of detailed and kinematics modelling level are plotted for

Figure 2. Simplified discretisation and modelling levels of the ‘frame�stringer�skin’ model.

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closing bending load, which leads to a compression-

loaded inner flange. Curve oscillations at increased strain

levels indicate the initiation of buckling at the compres-

sion-loaded inner flange. The bending stiffness of the

macro element in the kinematic hinge is calibrated so that

the strain distribution along the frame flange of the kine-

matics model and the detailed model agrees.

Figure 3(b) compares the strains of both modelling

levels for opening, bending and loading, which leads to a

tensile-loaded inner flange. Good agreement for both

modelling levels can be observed up to the initiation of

instabilities in the compression-loaded skin, which is indi-

cated by curve oscillations.

The plots of curves in Figure 3 show an abrupt

increase of the kinematic hinge strain (model L#4) which

is caused by the triggering to represent frame failure. The

failure initiation of the kinematic hinge leads to an abrupt

unloading of the adjacent frame structure with corre-

sponding decrease of the strains. In contrast to this, the

detailed model L#1 is defined with linear-elastic material

law wherefore failure is not represented here. The corre-

sponding trigger moments are determined by failure

criteria on the basis of strain values and are discussed in

the next paragraph.

Frame failure criterion

The kinematics model approach also comprises efficient

failure criteria to assess the structural loads and potential

exceeding of structural loading capacity. The longitudinal

strains at the frame flanges are used to quickly determine

the structural loading and to conclude potential failure.

The failure criteria on the basis of these longitudinal

strains can be defined according to handbook criteria that

consider for example the laminate strength for tensile

loading and crippling failure for compression loading

[6,13].

The omega-shaped frames used in this study are

equipped with comparably large mouseholes (Figure 2)

with the tendency of instability failure in the frame web.

Simplified handbook criteria are not adequate to cover

these failure modes. Hence, enhanced methods were intro-

duced, which consider the Hashin damage initiation crite-

ria using the detailed modelling level (L#1) of the

frame�stringer�skin model to derive the simplified fail-

ure criteria for longitudinal strains in the frame.

Pure bending loads are applied on the detailed fra-

me�stringer�skin model up to the exceeding of the

Hashin failure criteria at an arbitrary location of the struc-

ture. The Hashin criteria are checked by the contour plot

as illustrated in Figure 4. At this state of first exceeding of

the Hashin criteria, the average strain along the most criti-

cal inner frame flange is determined and defined as the

failure strain for this load case that can be used in the fuse-

lage section simulation for efficient analysis of the struc-

tural loading. Figure 4 illustrates this procedure for the

closing and opening bending load case.

Development of the crash kinematics

The crash kinematics of the tension crash concept was

developed on the basis of a statically pre-sized generic

single aisle full CFRP transport aircraft design that was

provided by Airbus in the scope of this research coopera-

tion. A two-bay fuselage section is considered under

purely vertical impact conditions.

The length of the fuselage section is 1270 mm with a

frame pitch of 635 mm. The frames are equipped with an

asymmetric omega-shaped cross section as depicted in

Figure 2. The fuselage cross section is described by four

different radii between 1887 and 2609 mm. The vertical

position of the cargo floor is about 300 mm (distance

between the lower skin and cargo floor level). All struc-

tural parts are modelled with shell elements, except the

stringers, sub-cargo and seat struts, which are modelled

with beam elements. According to the kinematics model

approach, elastic material formulations were used without

Figure 3. Calibration of the kinematic hinge macro element(frame longitudinal strains): (a) closing moment loading and (b)opening moment loading.

528 P. Schatrow and M. Waimer

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the definition of damage and failure, which is exclusively

represented by the macro models. An orthotropic material

model (Abaqus: �elastic, type ¼ lamina) was used for the

shell section definitions. Isotropic linear-elastic behaviour

was assigned to the beam elements. The elastic material

parameters are defined according to typical CFRP mate-

rial. The metallic parts, which are the seat rails, cargo rails

and seat structure, are defined with elastic�plastic mate-

rial formulations (Abaqus: �plastic, hardening ¼ isotro-

pic) based on aluminium Al2099 material data. Strain rate

dependent material behaviour was not considered in this

preliminary design study.

The total mass of the two-bay fuselage section is

1432 kg. The fuselage section model is equipped with two

seat rows of triple seats. Standard passenger masses of

77 kg were scaled to 68.5 kg to reasonably match the seat

pitch with the frame pitch in the simulation model. In

Table 1 the mass balance of the simulation model is

shown. An initial velocity of 6.7 m/s (22 ft/s) and a grav-

ity field of g ¼ 9.81 m/s2 in vertical direction was

assigned to all nodes of the fuselage section model. Sym-

metrical boundary conditions are applied on the front and

rear end of the simulation model.

As previously described, discrete rivets are not mod-

elled in the simulation model. Instead the structural parts

were tied to each other. A general contact between the

fuselage skin and the impact surface is defined with a fric-

tion coefficient of 0.4. In Table 2 the contact and tie part-

ners are shown. The passenger-seat system is modelled

using a mass element for passenger representation and

two-node connector elements for modelling the seat cush-

ion and seat back stiffness as well as the harness system.

The two-bay fuselage FE-model contains approxi-

mately 40,300 nodes and 40,000 elements with 32,500

shell elements, 5800 strain bar elements, 1700 beam ele-

ments, macro elements and mass elements. The total num-

ber of variables is approximately 220,000. The explicit

simulations were performed with the commercially avail-

able FE-code Abaqus/Explicit V6.11-1 on one CPU of a

Linux-Cluster. The simulation time is 200 ms, with an

Figure 4. Derivation of a simplified frame failure criterion based on frame longitudinal strains: (a) closing moment loading and (b)opening moment loading.

Table 1. Mass balance of the simulation model.

Structural element Mass

89% passengers (different frameand seat pitch)

12 � 68.5 kg ¼ 822 kg

Hatracks 2 � 75.8 kg ¼ 151.6 kg

Seats 131.7 kg

Systems 72.3 kg

Structure 254.4 kg

Two-bay simulation model 1432 kg

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initial time increment of 1e�03 ms, and led to a calcula-

tion time of about 11 hours.

In Figure 5 the approach for the crash scenario devel-

opment is illustrated.

In a first step detailed locations of the kinematic

hinges and potential trigger load levels for the individual

crash devices were determined by computing the simula-

tion model solely with linear-elastic properties. The strain

distribution along the frame was used to position kine-

matic hinges at locations of local strain extremum, which

indicate potential failure in the first phase of the crash

sequence. In addition, force output of the linear-elastically

modelled tension absorbers was analysed to determine

appropriate trigger load levels.

In a second step a reference setting was defined with

appropriate trigger load levels in the individual crash

devices and kinematic hinges that are positioned at poten-

tial failure locations of the frame.

In a third step a parameter study was performed to

analyse the influence of different parameters of the macro

element characteristic for tension absorption in the cargo

and cabin floor and for bending absorption in the frame.

Based on the results of the parameter study and outcomes

of several drop tests, which were performed in the past

[1,2,17,20,25,27�30,38,43,44] a general crash kinematics

was selected that aligns to the natural structural failure

behaviour of typical fuselage sections. In this crash kine-

matics, first frame failure occurs at the impact point fol-

lowed by subsequent frame failure directly below the

connection of the vertical support struts.

In Figure 6 the input characteristics of the macro ele-

ments are illustrated. Locations of maximum bending

loads in the frame are at the impact point in the kinematic

hinges (A) and below the vertical support strut connection

in the kinematic hinges (D). The crash sizing of the frame

structure (re-sizing of frame height and laminate thick-

ness) is achieved by measuring and assessing of the strain

along the inner frame flange. As a result no exceeding of

the failure strain should occur along the inner frame

flange, except in the kinematic hinges (A) and (D) that are

intended to fail.

In the kinematic hinges, the bending stiffness, Sf, the

opening trigger moment M1 and the closing trigger

moment M3 of the macro elements were adapted to the

surrounding frame structure as it was described previ-

ously. The post-failure absorption levels for opening

moment loading (M2) and for closing moment loading

(M4) are defined constantly at 30% of the corresponding

trigger moment and correspond to a low energy absorption

capability of the frame. Rigid behaviour is defined for the

remaining degrees of freedom of the macro element.

In the cabin floor, the tensile stiffness Su� of the ten-

sion absorber (d) is adapted to the tensile stiffness of the

passenger crossbeam, while the trigger force F1� corre-

sponds approximately to the ultimate load of the static siz-

ing. The constant absorption force level F2� is defined at

70% of the trigger force F1�. The ratio of 70% between

the steady state absorption force and the trigger force is a

typical value for bearing failure of composite laminates as

it can be observed from experiments performed in

[34,36,37]. The rotational stiffness of the cabin floor

absorbers were modelled linear-elastically about the x-

axis (flight direction). Rigid behaviour is defined for the

remaining degrees of freedom of the macro element.

In the cargo floor, the tensile stiffness Su of the tension

absorber (a) was assumed to be the same as that for the

cabin floor absorber. A detailed sizing of the cargo cross-

beam was not available in the preliminary design process

so that detailed stiffness values could not be derived from

the static sizing. The constant absorption force level F2

was chosen in this way that the remaining kinetic energy

that is not absorbed in the cabin floor tension absorbers

(d) and in the frame kinematic hinges (A) and (D) is

absorbed in the cargo floor tension absorbers. According

to this approach a significant amount of energy is

absorbed by tension loads. Finally, the trigger force F1

was determined with the relation F1 � 0:7 ¼ F2. The

macro element rotation of the cargo floor tension absorber

Table 2. Connection partners.

Structural parts Connection

Stringers to skin Tie

Frames to skin Tie

Sub-cargo structure to frames Tie

Hatracks to frames Tie

Seats to seat rails Tie

Seat rails to floor panels Tie

Seat rails to passenger beams Tie

Skin�impact surface General contact

Figure 5. Approach for the development of the tension crashkinematics.

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is free about the x-axis (flight direction). The remaining

degrees of freedom describe rigid behaviour.

The input characteristics of the macro elements are

defined without total rupture. The post-failure behaviour is

described by keeping a constant load level. With the maxi-

mum absorber displacements identified in the development

of this crash kinematics, physical absorber concepts can be

developed, which ensure energy absorption respectively

structural integrity up to the given displacements.

In this crash kinematics the vertical support struts do

not participate in the energy absorption; therefore, they

are modelled with linear-elastic input characteristics. The

rotation of the vertical support struts is fixed about the

x-axis (flight direction).

In Figure 7 the sequence of the crash kinematics is

shown. In the first crash phase (up to t ¼ 50 ms), the kine-

matic hinges (A) at the impact point fail, which causes the

triggering of the tension absorbers (a) in the cargo floor.

Subsequently, increasing crash loads and deformations

lead to the triggering of the kinematic hinges (D) and the

cabin floor tension absorbers (d) at approximately the

same time.

In the subsequent crash phase, kinetic energy is

absorbed in parallel by the frame bending mechanism of

the kinematic hinges (A) and (D) as well as by the tension

absorbers (a) in the cargo floor. In addition, some kinetic

energy is absorbed through ovalisation of the fuselage sec-

tion in the cabin floor tension absorbers (d).

Energy balance

The energy balance is described using general equations (1)

and (2):

ETOT � EI þ EKE þ EFD � EW ; ð1ÞEI � EE þ EDMD; ð2Þ

where ETOT is the total energy, EI is the internal energy, EKE

is the kinetic energy, EFD is the frictional dissipated energy,

EW is the external work of the gravity field, EE is the elastic

strain energy and EDMD is energy dissipated by damage.

Here, the numerical energies (e.g. artificial strain energy)

and work done by contact and constraint penalties are

Figure 6. Input characteristics for the macro elements representing structural failure.

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Figure 7. Crash kinematics of the fuselage section: (a) iso: t ¼ 50 ms; (b) iso: t ¼ 100 ms and (c) iso: t ¼ 150 ms.

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assumed to be negligibly small. Also, the inelastic dissipated

energy and the viscous dissipated energy are neglected,

which can be stated due to elastic material formulation used

in the simulation model.

At t0 ¼ 0 ms the total energy is equal to the initial

kinetic energy:

E0TOT ¼ E0

KE: ð3Þ

At t1 ¼ 200 ms the energy balance is represented by

E1TOT � E1

I þ E1KE þ E1

FD � E1W : ð4Þ

Since the total energy of the system is constant, com-

bining Equations (3) and (4) results in

E0TOT ¼ E1

TOT;

E0KE � E1

I þ E1KE þ E1

FD � E1W ;

E0KE þ E1

W � E1I þ E1

FD þ E1KE:

ð5Þ

According to Equation (5) the initial kinetic energy

and the external work (gravity field) are transformed in

the internal energy and in the frictional dissipated energy.

The frictional dissipated energy is mainly related to slight-

ing friction between fuselage section and impact surface.

In Figure 8(a) the energy balance of the simulation

model is shown. The total energy ETOT is constant at 32.1

kJ. At the final crash state ðt1 ¼ 200msÞ, the initial kineticenergy and the external work E0

KE þ E1W (100%) are trans-

formed to the internal energy E1I (84.8%) and to the fric-

tional dissipated energy E1FD (7.2%). The remaining

kinetic energy E1KE (6.2%) is in the fuselage structure at

this final state of the simulation. Furthermore, neglected

energies (1.8%) as described previously are in the system

and represent the flaw size of Equation (5).

Noticeable is the smooth decrease of the kinetic

energy during the whole crash sequence, which is equiva-

lent to minimum acceleration loadings of the passengers

and minimum crash loads for the fuselage structure.

In Figure 8(b) the absorbed energies of the individual

crash devices are shown. With respect to the initial kinetic

energy and the external work ðE0KE þ E1

W Þ the amount of

the absorbed energy in the cargo floor tension absorbers is

36.4% (14.1 kJ) and of the cabin floor tension absorbers

8.0% (3.1 kJ). The amount of the absorbed energy in the

kinematic hinges (A) is 6.8% (2.6 kJ) and in the kinematic

hinges (D) 11.9% (4.6 kJ). Altogether about 63.1% (24.5

kJ) of (E0KE þ E1

W ) was absorbed by the macro models of

tension absorption and frame failure. The internal energy

(63.1%) of the macro elements consists of the elastic

strain energy (6.0%) and of the energy dissipated by dam-

age (57.1%). The amount of elastic strain energy in the

linear-elastically modelled fuselage structure, except the

macro elements, is 21.0%. The elastic strain energy of the

fuselage section model and the energy dissipated by dam-

age are not shown in the diagram.

As can be seen from Figure 8(b) the energy absorption

occurs simultaneously in all crash devices. This is a main

advantage of the considered crash concept as the parallel

interaction of different absorption mechanisms allows dif-

ferent design concepts with individual adaptation of local

absorption capacities. Furthermore, undesired load peaks

caused by triggering of the next level in a cascading crash

concept do not occur in this crash concept.

Output of the macro elements (displacement

and rotation)

Output data of the macro elements can be used to assess

the requirements given to the crash devices to obtain the

considered crash kinematics. In the following, output is

given with respect to one of the symmetrically installed

Figure 8. Energy output: (a) energy balance and (b) internalenergy in the crash devices.

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macro elements of each macro type as the crash sequence

provides almost purely symmetrical kinematics. In

Figure 9(a) displacement�time plots of the tension

absorber in the cargo floor (a) and the cabin floor (d) are

illustrated. Again, the tension absorber (a) shows a smooth

displacement curve resulting in constant energy absorption

during the crash sequence. The total tension displacement

in each of the two macro elements in the cargo floor is

121 mm. The tension absorber (d) in the cabin floor shows

a first activation in the initial crash phase up to t ¼ 30 ms.

Further tension displacement occurred during the ground

contact of the vertical support struts at about t ¼ 105 ms as

slightly increased total crash loads lead to further ovalisa-

tion of the fuselage. The total ovalisation displacement or

tension absorption on each side of the cabin floor is 43 mm.

The maximum unfiltered loading rate, which occurred

in the cargo floor tension absorbers is 6 m/s, in the cabin

floor absorbers the loading rate reached about 3 m/s �with an output frequency of 20 kHz.

Frame bending rotation of the kinematic hinges (A)

and (D) are given in Figure 9(b). According to the crash

kinematics, a plot of smooth rotation curves can be seen

resulting in constant energy absorption by frame bending

during the crash sequence. Positive and negative rotations

are equivalent to opening, respectively, closing bending

moments.

The total bending rotation in the kinematic hinges (A)

is 55.6� (0.97 rad), each of the kinematic hinges (D)

located close to the vertical support strut reached a closing

bending rotation of �33.2� (�0.58 rad).

Assessment of the structural loads in the frame

In the scope of this simulation study the focus of the struc-

tural loads was on the frame, which is the most important

crash relevant structure. Loads on the frame structure

were assessed using the above discussed longitudinal

strains, which are determined by bar elements of negligi-

ble stiffness positioned at the frame flanges.

The longitudinal strains were checked with respect to

the given failure criteria and the statically pre-sized frame

structure was re-sized according to the most critical

strains in the inner frame flange. All results of the fuselage

section simulation model given in this paper are with

respect to the final crashworthy frame sizing.

Figure 10 depicts plots of all longitudinal strains,

which are determined by the bar elements along the

frame. The crashworthy frame design provides a distribu-

tion of stiffness along the frame according to the crash

loads in the individual regions. This frame profile distribu-

tion also distinguishes in different failure strains for ten-

sion and compression, respectively, opening and closing

bending, which is observable in Figure 10 in a bandwidth

of different minimum and maximum failure strains.

Curves of bar elements for strain measurement, which are

located directly in kinematic hinges indicate by abrupt

increase the triggering of the kinematic hinges. The kine-

matic hinge (A) triggers immediately after the first impact

and the kinematic hinges (D) trigger at about 6 ms after

the first impact.

Few strain curves clearly exceed the failure strain val-

ues. These strains are influenced by rigid bodies in the

kinematic hinges and at the passenger beam connection.

Based on the selected modelling approach these artifi-

cially high strain values were not considered in the assess-

ment of the structural loads.

In addition, Figure 10 shows state diagrams, which

plot the strain distribution along the frame of one state in

a circle diagram against the circumferential angle of the

fuselage section. Critical frame regions in which failure

strain limits are exceeded can be visualised with these

plots. The individual failure strain limits of all frame pro-

files along the frame are plotted in the diagrams. Strains,

which represent kinematic hinges, are marked with a

white diamond. As can be seen from the state diagrams no

significant exceeding of the strain limits occur due to the

Figure 9. Displacement and rotation output data: (a) tensionmacro elements in the sub-cargo area and in the cabin floor and(b) kinematic hinges (A) and (D) in the frame.

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adaptation of the frame design, except some strains in the

upper fuselage section at t ¼ 70 ms and close to the pas-

senger crossbeam connection at t ¼ 100 ms. Despite these

strain exceedings, further adaptation of the crashworthy

frame design was not considered in this preliminary design

study as further crash load cases have to be regarded first

before a more detailed crash sizing is reasonable.

Assessment of structural loads in the cargo

and cabin floor

The cargo floor framework structure essentially contrib-

utes to the tension crash concept as its structural integrity

has to remain intact during the crash event to maintain the

bending mechanism that leads to tension forces in the

cargo crossbeam. Structural loads in the framework struts

have to be checked to assess the feasibility of a sub-cargo

structure to provide the required stiffness and strength.

In Figure 11(a) the longitudinal forces in the vertical

struts of the cargo floor framework are illustrated. The

unfiltered output data were sampled with a frequency of

20 kHz and afterwards filtered with a Butterworth filter

and a cut-off frequency of 0.1 kHz. The maximum unfil-

tered force in compression direction of the inner struts (1)

is �12.5 kN and in tension direction þ9.5 kN, while

the maximum filtered force in compression direction is

�11.5 kN. The maximum unfiltered force in compression

direction of the outer struts (2) is �20.0 kN and in tension

direction þ2.5 kN, while the maximum filtered value in

compression direction is �9.5 kN. The unfiltered maxi-

mum peak forces in the inner struts (1) and in the outer

struts (2) are caused by different effects. In the inner struts

(1) the maximum peak force occurs due to the dynamic

shock induced by the first impact on the ground surface.

In the outer struts (2) the maximum peak force is related

to the activation of the horizontal tension absorbers (a) in

the cargo floor (Figure 11(b)). The plot (Figure 11(a)) of

filtered strut forces shows increasing compression forces

in the inner strut (1) according to the crash kinematics, as

the tension absorber forces (a) provide an increasing com-

ponent in compression direction of strut (1) during the

crash sequence. This geometric relation can be seen in

Figure 10. Circumferential strain in the inner flange of the frame (state plots and overall plot).

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Figure 7. Oblique strut positions can reduce these com-

pression forces in the cargo floor framework struts with

respect to the considered crash kinematics.

The (horizontal) tension absorber forces in the cargo

and the cabin floor are shown in Figure 11(b). The cargo

floor tension absorbers trigger at a comparably high load

level of 42.5 kN. Although a reduced trigger force can be

realised easily, this high load level illustrates the potential

to absorb energy by tension mechanisms due to the high

loads that can act in the cargo crossbeam. The constant

force level of the tension absorber (a) in Figure 11(b) rep-

resents smooth energy absorption respectively displace-

ment as shown in Figure 9(a). In contrast to this, the cabin

floor tension absorber (d) shows in Figure 11(b) a discon-

tinuous force curve that can be compared with Figure 9(a)

to understand the unsteady absorption according to the

ovalisation behaviour of the fuselage section. The cabin

floor tension absorber is triggered at a load level of 25 kN,

which corresponds approximately to the ultimate load of

the passenger crossbeam, and therefore the minimum trig-

ger load for controlled failure in case of crash.

Assessment of the passenger loads

In the simulation model the passengers are represented by

single mass elements, which are connected to the seat

structure by spring elements that represent the seat cush-

ion stiffness and the harness system [40]. Non-linear

force-displacement characteristics as well as hysteresis

behaviour of the seat base cushion were calibrated on the

basis of available test data (pelvis-cushion compression

test) [26]. According to this modelling approach, the

acceleration response of the passengers were assessed in

an acceleration-time diagram and in an Eiband diagram

[16,23,24].

In Figure 12(a) vertical acceleration responses are

illustrated in an acceleration�time diagram. The passen-

gers considered here are located in the first seat row on

the left side of the fuselage section and represent an over-

view on the passenger loads due to the tendency of struc-

tural symmetry in the simulation model. The responses

are unfiltered and are sampled with an output frequency

of 20 kHz. The passenger at the aisle (PAX 1-C) experien-

ces the largest peak acceleration of 41.5g. The passenger

next to the window (PAX 1-A) experiences a peak accel-

eration of 34g. The lowest peak acceleration of 24.5g is

acting on the passenger in the middle seat (PAX 1-B).

With respect to this distribution of passenger loads in a

seat row, the cabin floor dynamic behaviour is expected to

highly influence the acceleration responses.

In Figure 12(b) the vertical acceleration responses of

the considered passengers are plotted in the Eiband dia-

gram. Limit curves of moderate and severe injury for

headward acceleration according to Eiband are given in

this diagram. The acceleration responses of the passenger

at the aisle (PAX 1-C) and next to the window (PAX 1-A)

are in the upper region of moderate injury but still distant

from the limit of severe injury. The acceleration response

of the passenger in the middle seat (PAX 1-B) is next to

the limit of moderate injury. Finally, all passenger loads

are below the limit of severe injury.

Summary and conclusion

Special crash devices are required in a CFRP transport air-

craft to absorb the kinetic energy similar to equivalent

metallic fuselage structures and to limit the passenger

loads in survivable crash accidents. Typical absorption

devices, which are considered in this context, are based

on crushing failure mechanisms in the struts of the cargo

floor framework or on bending mechanisms in the frame.

The purpose of the present study is to show alternatives in

the absorption of kinetic energy using tension mechanisms

in crash devices, which are positioned in the cargo floor

and cabin floor. These absorbers benefit from the tension

loads acting in the cargo crossbeam and the passenger

crossbeam in typical crash events. A crash kinematics

Figure 11. Force output: (a) longitudinal forces in the verticalstruts of the sub-cargo framework and (b) longitudinal forces inthe tension absorbers of the sub-cargo area and the cabin floor.

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with tension absorption as the main energy absorption

mechanism was investigated using an FE simulation

model of a two-bay fuselage section, in which the crash

devices are represented by macro elements with force-dis-

placement and moment-rotation input characteristics. The

remaining structure, that is intended to experience no

damage, is modelled linear-elastically. The developed

crash kinematics is characterised by frame failure at the

impact point and below the vertical support strut connec-

tion that leads to an unrolling kinematics of the lower

fuselage section with tension absorption mainly in the

cargo floor.

After development and assessment of the crash con-

cept with tension absorption in the cargo floor the follow-

ing outcomes can be stated:

� Tension absorption mechanisms generally require a

less complex structural environment due to the pre-

vention of instabilities, resulting in less massive back-

ing structures compared to compression absorbers.

� The developed tension crash kinematics allows paral-

lel energy absorption in several crash devices, e.g. the

cargo floor tension absorbers and the frame bending

mechanisms (kinematic hinges). Different absorption

characteristics can be combined flexibly leading to an

appropriate total energy absorption as required.

� Parallel activation of the main absorption mecha-

nisms prevents load peaks which can occur from

cascading crash concepts by successive triggering

of the individual absorbers.

� The developed crash kinematics with tension

absorption shows almost constant energy absorption

during the whole crash sequence with smooth

decrease of kinetic energy.

� The crashworthy frame sizing conducted in the

scope of this study identified the tendency to a

reduced structural mass penalty compared to the

bend frame crash concept that specifies crushing

failure of the sub-cargo structure.

� Finally, the simulation study identified potential

absorber characteristics that are necessary to

achieve the desired tension crash kinematics.

Required load levels for failure initiation (trigger

loads), absorption load levels and maximum

absorber displacements can be used to develop

appropriate tension absorber concepts.

Acknowledgements

The authors wish to thank Dieter Hachenberg and Lars Margull(Airbus) as well as Brian Bautz, Tim Bergmann and SebastianHeimbs (Airbus Group Innovations) for good support providedby technical discussions and information about structuraldesigns. The research was conducted in close collaboration withAirbus Operations GmbH and Airbus Group Innovations.

Funding

The research leading to these results was accomplished in theframework of the 4th Aeronautical Research Programme of theGerman Federal Ministry of Economics and Technology(BMWi) [grant number 20W1105B] as part of the LuFo-IV proj-ect CREVAD (TENOR).

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