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Formatted as in InternationalJournal of Impact Engineering 35 (2008) 1075

Oblique impact testing of bicycle helmets

N.J. Mills* and A. Gilchrist

Metallurgy and Materials, University of Birmingham, Birmingham B15 2TT, UK.

Received 25 November 2005; received in revised form 5 February 2007, accepted 24 May 2007

*corresponding author E-mail address [email protected]

Abstract

The performance of bicycle helmets was investigated in oblique impacts with a simulated road surface. The linear and rotational accelerations of a headform, fitted with a compliant scalp and a wig, were measured. The peak rotational accelerations, the order of 5 krad s-2 when the tangential velocity component was 4 m s-1, were only slightly greater than in comparable direct impact tests. Oblique impact tests were possible on the front lower edge of the helmet, a site commonly struck in crashes, without the headform striking the ‘road’. Data characterizing the frictional response at the road/shell and helmet/head interfaces, was generated for interpretation via FEA modelling.

Keywords: bicycle, helmet, impact, foam, protection ______________________________________________________________________________________

1. Introduction

Accident analysis shows [1] that the majority of bicycle helmet impacts are oblique to a surface, with the most common impact sites at the front and sides of the head [2]. Hodgson [3] showed that high frictional forces could occur when no-shell bicycle helmets made oblique impacts on rough concrete surfaces, while McIntosh [4] noted that some no-shell helmets fractured into several pieces in crashes. Consequently, current helmets all have an external shell. This paper describes bicycle helmet performance in oblique impact tests; a parallel paper [5] uses some results to validate finite element analysis (FEA) of the tests. In return, FEA can interpret experimental events, reveal causative mechanisms, and allow the evaluation of interface frictional conditions.

The first aim was to construct an instrumented test rig, to measure the key parameters when a helmeted headform makes an oblique impact with a road. Aldman et al. [6] dropped a complete dummy, held horizontally, onto a rotating turntable of diameter 1 metre. They found, for tests on motorcycle helmets, that the peak headform rotational acceleration was approximately independent of the tangential velocity component, in the range 0 to 11 m s-1. However, their test rig was large and dangerous, while the use of a whole dummy was not necessary. Tests with a dummy with a realistically flexible neck [7] showed that the head motion in the first 40 ms of an impact is

unaffected by the neck; the impact force on the helmet was equal to the linear acceleration of the head multiplied by the headform mass. Studies of the load transfer through the necks of vertically-dropped cadavers [8] confirm this conclusion. Consequently, the peak headform rotational acceleration in an oblique impact test using a free headform should be the same as when using a full dummy. In a ‘2-Dimensional’ experiment, a 25 mm thick slice of a bicycle helmet, bonded to a constant-radius ‘headform’ [9] made an oblique impact on a surface covered with SiC paper; the coefficient of friction of the helmet shell was measured as about 0.3. Rather than drop a headform plus bicycle helmet vertically onto a fixed inclined rough plane representing a road, it was preferred to move the ‘road’ horizontally and use a lower drop distance [10]. The test rig initially had no force cells, and only a single rotational accelerometer in the headform; its development is described here.

A second aim was to gather data to validate FEA of helmet oblique impacts. It is impractical to instrument helmets with strain gauges, since gauges mounted on the shell exterior would interfere with the impact, and gauges cannot be successfully mounted on the expanded polystyrene bead foam (EPS) liner. However, by using biaxial load cells in the road surface, rotational acceleration transducers in the headform, and high-speed film, the headform rotation can be evaluated.


The third aim was to evaluate the effects of the helmet external shape, and the impact site, on the peak headform rotational accelerations, which relate to the risk of brain injury. To obtain meaningful results, the helmet must rotate realistically on the headform. Hence details of the helmet retention system are important, while scalp and hair must be simulated on the headform. Skin is anisotropic and the underlying muscle and adipose tissue allow easy shear of the skin relative to the skull, as noted by palpation. The scalp thickness averages about 5.5 mm with an adipose layer of 3.1 mm [11]. There is no published shear stiffness data for the scalp, while adipose tissue has a shear modulus the order of 5 kPa [12].

The final aim was to investigate assertions made by anti-helmet campaigners. Curnow [13] argued that bicycle helmet design reflects a discredited theory of brain injury (that injuries are caused by peak linear acceleration). He correctly stated that bicycle helmet standards do not contain oblique impact tests in which the headform rotational acceleration is measured. However his premise, that the majority of bicyclists’ head injuries are due to rotational acceleration, can be refuted by statistics of the type of head injuries suffered. Franklin [14] also criticised the lack of oblique impact tests in helmet standards, writing ‘there does not appear to be research evidence that cycle helmets are effective in mitigating angular impacts.’ Henderson [15] was critical of vertical drop tests, noting that ‘the solid headform used for standards approval does not mimic the deformable characteristics of the human head’.

2. Bicycle helmet design

Bicycle helmet design has changed markedly since 1990 [16]. When British Standard BS 6863 was replaced by EN 1078 [17] in 1997, the impact test drop height increased from 1.0 to 1.5 metres. Current helmets have more ventilation holes, are thicker at the rear, and sometimes have a non-smooth external profile. The helmets subjected to oblique impact tests in 2002 [10] had less than ten, large ventilation holes. By 2005 the number of ventilation holes has increased, and their size decreased. The direction of the holes is affected by the cost of mould design; cheaper helmets tend to have two-part moulds with a single opening direction (vertical on the helmet as worn). More expensive moulds can split into multiple sections, so holes at the front and rear of the helmet can be in the direction of travel. Figure 1 compares ventilation holes in a Specialized S1 helmet sold for circa £100, and in an Aventicum helmet sold for £6 by Aldi. Figure 2 shows side views of two current designs. Projections at the rear possibly aid airflow through the helmet, while the outer surface of the shell is relatively smooth.

The majority of helmets have adjustable-circumference headbands (Fig. 3b), with a single size of shell and liner, while the Specialised S1 helmet has a patented [18]

moulding that grips the nape, via a spring-loaded mount at rear (Fig. 3a). Helmets are supplied with small soft foam pads, attached by Velcro at several sites inside the liner, to bridge the gap with the head.

3. Helmets tested

Helmets, certified to EN 1078, were chosen to represent a range of designs (Table 1). The dates of manufacture were between August 2004 and March 2005. As helmet positioning on the headform might affect the impact results, a standardized position was used. The helmets were fitted to an Ogle headform, which is 192 long and 155 mm wide at the level of the AA' plane - a level referred to in EN960 [19] as the headband level. For an 570 mm circumference headform, the basic plane (defined as passing through the external ear opening and the lower edge of the eye orbits) is 40 mm below the AA' plane, and 90 mm below crown of the headform. The helmet front brim was adjusted to be 2 cm above the AA' plane, while the chin strap was adjusted to be tight below the chin. The helmet liner thickness at specific sites falls within a narrow band. The sites are defined by the rotation of the head and helmet from the crown site (figure 4a), so the right 70° site will be at the lowest part of the helmet after a 70° rotation about the 3 axis. All the liners taper in thickness above the ear, so the value on the right 70° site is an average with a ± 3 mm error.

Table 2 gives details of the retention systems. Headbands fixed to the liner by moulded plastic inserts in the foam are more secure than those fixed using Velcro. The majority of front webbing straps pass through holes in the EPS, then between the EPS exterior and the shell interior, being either fixed to the liner by adhesive, or free to slip. The S1 helmet side straps are attached to moulded inserts inside the EPS liner. The lower the attachment points (relative to the lower edge of the helmet) and the further apart laterally, the more effective is the strap geometry in resisting helmet rotation. Thus Table 2 probably ranks the helmets in decreasing order of strap location effectiveness. For all helmets bar the cheap Aventicum, the rear webbing straps were fixed to the liner centrally at the rear, about 6 cm above the lower edge. In the latter the strap was looped through holes in the liner, 13 cm above the lower edge.

4. Oblique impact test rig

A free-falling headform impacted a horizontally-moving sandwich, consisting of two aluminium plates (Fig. 4b) 0.5 m long, 0.15 m wide and 10 mm thick separated by two triaxial quartz force cells (Kistler 9348B). These very rigid cells are each bolted at all four corners (40 mm separation) to both plates, creating a stiff structure. The signal charges were added before amplification, so the total


Fig. 1. Plan view of vents in helmets, with front at top: a) Specialized S1, b) Aventicum

Fig. 2. Side view of helmets with front at right: a) Specialized S1, b) Giro Indicator (note the cracks in the foam after testing)

normal FN and tangential FT forces on the top plate were measured. Vibration-damping layers (Heathcote Industrial Plastics, type 4005) were used to reduce plate vibrations. The FN versus time response of the non-moving sandwich for a vertical helmet drop was the same as for an impact into a fixed rigid load cell, showing vibrations of the aluminium plate have no effect on the response. The assembly, of total mass 7.458 kg, is moved by a pneumatic cylinder of 1 m stroke while resting on flat, horizontally supported, PTFE bearings. For some experiments, a rough road surface was simulated by 120 grade SiC grit grinding paper, bonded to the upper plate.

However, for most experiments an aluminium surface, abraded with 80 grade SiC paper, was used. The helmet position was maintained during its vertical fall by a U-shaped aluminium honeycomb frame, attached via a 4-wheeled carriage to a vertical monorail. The frame continues downwards, parting from the helmet, when the latter impacts the horizontal plate. The pre-impact vertical velocity of the helmet was measured between photocells, 6 and 1 cm above the impact plane. It was possible to vary the initial helmet position, the helmet vertical velocity component up to 6 m s-1 and the horizontal plate velocity up to 4 m s-1. The majority of helmets were dropped from 1.5 m, as in EN 1078, but friction in the monorail bearing


Fig. 3. Rear view of helmets, showing the size adjustment mechanisms: a) Specialised S1, b) Giro Indicator. Table 1. Details of helmets tested Manufacturer model

Giro Indicator

Bell Avanti

Bell Arc

Grepper Aventicum

Specialised S1

Size universal universal universal large large Liner density kg m-3 81 81 90 71 101 Total Mass g 260 321 321 270 247 Shell mass g 25 35 30 40 inseparable Shell bonded to liner yes yes taped at rim + taped at rim yes Shell thickness mm 0.5 ± 0.05 0.4 ± 0.05 0.35 ± 0.05 0.5 ± 0.05 0.6 ± 0.1 length inside at base mm 220 220 220 230 220 width inside at base mm 175 180 175 180 175 Liner thickness mm

Front 90° 29 27 28 24 22 Right 70° 24 22 23 21 23

Crown 28 30 26 28 26

meant that the impact velocity was less than the 5.42 m s-1 specified in the standard.

An internal vibration-damping layer was added to a hollow aluminium headform (Ogle Ltd) with an external PVC plastisol skin (figure 5). Its 4.26 kg mass is slightly less than the 4.7 ±0.14 kg specified in EN 960 [19] for circumference 570 mm. At its centre of gravity there was a Kistler 8792 quartz triaxial linear accelerometer and two Kistler 8838 rotational accelerometers, one aligned with the neck to crown (z) axis, and the other either on the ear to ear (y) axis or the nose to occiput (x) axis. The PVC plastisol skin on the rear, 50 mm deep, removable section of the headform was replaced with a 4 mm thick layer of Astrosorb M3 soft rubber (www.astron2000.at) to reduce the circumference – this area was not involved in impacts, but the change probably affected the ease of helmet rotation. The shear modulus of the main PVC scalp layer was measured as 73 kPa, an order of magnitude match for that of the human scalp. Drops of the headform from 10 cm onto a flat rigid anvil [10] showed that its contact stiffness was comparable with Allsop et al’s 6.9 Nmm-1 [20] for an

impact on the temporal-parietal region of a shaved cadaver head.

The headform was covered with an acrylic wig, which potentially provides a slippery interface with the helmet. The headform circumference is 580 mm over the wig. The moments of inertia of the headform, measured with a torsion pendulum, are compared with data for the 70th percentile of the male population [21] in Table 3. Moments of inertia of cadaver heads were measured on a torsion pendulum at a frequency of about 1 Hz [22], during which the soft tissues do not deform. The values at impact strain rates are likely to be lower, due to the rotation of the brain, floating in cerebrospinal fluid, inside the skull. The headform axes are shown in figure 4a. However van der Bosch [22] used axes rotated 30° in the sagittal (xz) plane, so their x′ axis is 30° up from the x axis, while z′ is 30° back from the z axis. The Ogle headform, with a partial neck, has a larger Iyy than Van der Bosch’s head-alone value; their value increases to 371 kg cm2 with a neck.


Fig. 4. a) Headform and instrumented plate axes, and velocity components. The crown impact site is shown. b) the ‘road’ instrumentation, with load cell centres 0.24 m apart

Table 2. Headband and retention strap details model headband

adjustment Headband

fixing to liner Front strap locations through liner

(cm) up lateral

Front strap

S1 no headband 0 ±8.2 fixed in liner Indicator ratchets 3 inserts & rear hanger 4 ±5 fixed to moulding at liner surface Avanti wheel 4 inserts 6 ±4.5 fixed to liner surface

Arc ratchets 2 inserts & rear hanger 9 ±3.3 between liner +shell Aventicum wheel Velcro to liner 10 ±4 free between liner & shell

The helmet chin straps were securely fastened over a 10

mm thick layer of compliant Airex S50 foam under the headform metal chin, simulating typical strap tightness.

The acceleration traces were recorded with 12 bit accuracy at 5 kHz without filtering. However the rotational acceleration signal can be noisy during and after the linear acceleration peak. Since rotational acceleration peaks lasting < 1 ms are unlikely to be injurious, the signals were filtered by a fast Fourier transform method with a cut-off frequency of 2000 Hz.

A linear accelerometer, with axis aligned with the direction of motion of the road surface, was mounted on the side of one load cell (Fig. 4b). However, its output a suffers from drift during the 250 ms of piston motion before the impact. Hence computation of the pre-impact road velocity from this signal would be inaccurate. The load cell signal does not drift with time, but acceleration of the road

surface, prior to impact, causes an apparent signal RTF .

Comparison of the two signals during the pre-impact acceleration of the road surface shows that

amF PR

T= (1)

where the effective mass of the upper plate and the upper parts of the load cells mP = 3.726 kg. Integration of R

TF / mP with respect to time gives the horizontal pre-impact velocity VH. The tangential force on the helmet during the impact is calculated using

amFF PR

TT −= (2)

Thus, FT is zero prior to the helmet impact.


5. Results

5.1. Rotational acceleration in oblique impacts

When the side of a helmet makes an oblique impact, it is most likely that the road moves rearwards relative to the helmet. The lowest likely impact site without the rider’s shoulder hitting the road was chosen: for these left 70° and right 70° impact sites, the helmet is expected to rotate about the neck-to-crown z axis (Fig. 6) relative to the head. The oval shape of the head and helmet horizontal cross-sections eventually limit the rotation. If a cyclist tumbles forwards over the handlebars, a likely impact site is towards the front of the helmet, with the road surface moving downwards relative to the face. The helmet liner is expected to rotate on the head about the ear-to-ear y axis.

Fig. 5. Instrumentation inside the Ogle headform, seen when the rear section is removed.

The rotation is resisted by the retention-strap system and limited by the liner front interacting with the nose. EN1078 standard impacts, using a metal headform falling vertically onto a fixed flat anvil, are not performed at a ‘front 90°’ impact site (achieved by rotating the head and helmet from the crown impact position by 90° about the 1 axis), probably because the headform would rotate until its face struck the anvil, damaging the headform and its instrumentation. However, in an oblique impact, the front 90° site can be used, as frictional forces on the helmet resist headform rotation, and can prevent face-to-road contact – figure 7 shows frames from a high speed film of such an impact.

The acceleration and force traces vs. time for a typical helmet (Figure 8) have a single peak with some superimposed oscillations; Table 4 gives the peak values for the tests. For the frontal 90° impact site, the peak headform linear accelerations are less than the 250 g limit set in EN 1078. The peak headform rotational accelerations are typically 5 krad s-2. The headform rotational velocity components after oblique impacts on the side 70° sites were a maximum of 25 rad s-1 and usually about 15 rad s-1, whereas they were less than 5 rad s-1 for impacts on the front 90° site.

Figure 9 shows how the magnitude of the net head rotational acceleration in the 12 plane varies as a function of the net headform linear acceleration; if FEA can predict this graph, the frictional parameters at the head/helmet interface can be estimated [5].

5.2 Rotational acceleration in direct impacts

For direct impacts (with the road surface stationary) on some test sites, the headform was seen to rotate on rebound, indicating rotational acceleration during the impact. Table 5 shows the values measured. For a crown impact site, the z axis rotational acceleration changed sign about halfway through the impact, while the y axis rotational acceleration component remained small.

Tangential forces during oblique impacts

A typical tangential force FT vs. time trace (Fig. 8b) has a single peak with some superimposed oscillations. The normal force FN vs. time trace also has a single peak. A plot of FT against FN, for many tests, has a trend line slope of approximately 0.2 (Table 4, with a typical trace shown in Fig. 10). These graphs will be interpreted in [5].

Table 3. Headform dimensions, mass and inertia Headform Circumference

mm Length

mm Breadth

mm Mass

kg Ixx

kg cm2 Iyy

kg cm2 Izz

kg cm2 Ogle test 580 190 158 4.26 199 237 172 used in FEA 199 154 4.26 192 240 163 Van der Bosch[24] 600 4.4 206* 178 151*


0 ms 10 ms

20 ms 30 ms Fig. 6. Frames of high speed video showing the Aventicum helmet rotating after an impact on the right 70° impact site.

5.4. Headform linear acceleration as a function of liner crushing

In direct impact tests, as the motion of the headform centre of gravity is approximately along a straight line, it is possible, knowing the initial impact velocity, to integrate the headform acceleration twice with respect to time [16] and calculate the helmet liner crushing distance. The maximum value, as a percentage of the liner thickness, estimates the amount of helmet protection used in the impact, a quantity useful for design purposes. The headform angular position in an oblique impact test can only be determined by analysing the output of a complex array of accelerometers [23]; this angular position is, in general, needed to determine the lowest position of the headform surface. However, for the impacts described here, the headform rotations are relatively small during the contact phase, and the largest linear acceleration component is normal to the road surface. Hence,

calculations of the maximum helmet liner crush can be made for oblique impacts. FEA [5] was used to validate these calculations. Using impact velocity components and sites typical of experiments, the magnitude of the headform linear acceleration vector a was predicted as a function of time. The vertical component a2 was put equal to a, then numerically integrated twice to estimate the vertical position x2 of the headform CG. As the headform surface is locally nearly spherical, a small headform rotation hardly changes its ‘vertical’ radius, hence the liner thickness. Consequently, an increase in x2 should produce an equal increase in the foam liner compression xL in the centre of the contact region. The estimated xL is compared with the value measured from a 23 plane projection of the headform and helmet, as a function of headform impact force (the product of its mass mH and acceleration magnitude a2) in Figure 11. xL is accurately estimated during the loading phase, but is overestimated by 2 or 3 mm just after the load maximum, when the helmet is rolling on the road.


0 ms 10 ms

20 ms 30 ms

Fig. 7. Frames of video of oblique impact with VN = 4.5 m s-1 and VT = 3.6 m s-1, on frontal 90° site of an Arc helmet.

.

Therefore, the loading slopes, measured before the maximum load, should be accurate, whereas the maximum deformations could be underestimated by 2 or 3 mm. For the frontal 90° impacts, in which there is less headform rotation, the errors should be less. During the initial 5 mm of deformation, the liner compresses the wig and comfort foam until it contacts the headform at the impact site, and the headform force is low. For most impacts, the subsequent portion of the graph has a

near-linear increase (the ‘loading slope’). However, some traces show two or three large superimposed oscillations (the result marked ‘oscill’ in Table 4) while one result in Table 4 had a two-stage response, with a low slope being succeeded by a higher slope. The headform force then reaches a peak, before dropping to near-zero, with very little liner thickness recovery on unloading. Tables 4 and 5 give the maximum liner crush distances, and the loading slopes. In general, for normal velocity components of


-2

-1

0

1

2

3

4

5

-50

0

50

100

150

0 2 4 6 8 10 12

rota

tiona

l acc

eler

atio

n

krad

s-2

time ms

linea

r acc

eler

atio

n g

y axis

a

z axis

-1

0

1

2

3

4

5

6

0 5 10 15

forc

es o

n ro

ad s

urfa

ce

kN

time ms

FN

FT

a) b)

Fig. 8. Oblique impact with VN = 4.5 m s-1 and VT = 3.6 m s-1 on left 70° site of Aventicum helmet: a) resultant linear and rotational acceleration components vs. time, b) force components on road surface vs. time Table 4. Oblique impacts with tangential velocity 3.6 m s-1 and normal velocity 4.5 m s-1 helmet Impact

site Max. Accel.

G

Max. FN kN

Max. FT kN

Max. z axisrot. acc. krad s-2

Max. y axisrot. acc. krad s-2

max. liner crush mm

Loading slope N mm-1

FT FN

Aventicum 114 5.2 0.5 3.9 3.6 17 320 0.21 Arc 129 6.0 0.8 4.0 4.9 15 420 Avanti 121 5.6 0.7 3.7 3.0 16 340 0.25 Indicator

Left 70°

115 5.5 0.7 3.9 2.8 16 340 0.23 Indicator * right 70° 109 5.1 0.5 -4.3 6.2 19 250 Specialized Left 70° 129 5.9 0.7 4.7 5.6 13 490 0.18 Avanti 106 4.8 0.9 < 1.0 & > -1.0 19 230 Aventicum 105 4.8 0.4 < 1.5 & > -1.5 15 340 0.20 Arc

Front 90°

117 5.5 0.5 < 1.0 & > -1.0 Two stage 0.22 *The impact surface was 100 grade SiC paper

Fig. 9. Oblique impact with VN = 4.5 m s-1 on right 70° site of S1 helmet: net rotational acceleration vs. net head linear acceleration

.4.5 m s-1 onto a flat surface, the foam liners are about 75% crushed at the right 70° impact site. Figure 11, which confirms earlier [16] findings of a near-linear increase in the impact force with the crush distance, contradicts Burdett’s [24] assertion that there is a minimum force before the helmet deforms. The maximum linear headform acceleration is slightly reduced by the tangential velocity component; the Arc (or Avanti) helmet directly impacted on left 70° sites at 4.5 m s-1 had a 138 (135) g peak acceleration, compared with 129 (121) g when a 3.6 m s-1 tangential velocity component was added at the equivalent site. FEA [5] predicted a similar slight decrease in the peak headform linear acceleration between VH = 0 and VH = 5 ms-1.

5.5 Foam fractures and helmet rotation

In many of the helmets, oblique impacts caused one or more liner fractures, either in lower regions not covered by the shell (Fig. 2b) or beneath the shell, but the microshell remained intact. The sliding distance was only a few mm, judging from the length of scratches on


Table 5. Direct impacts with normal velocity 4.5 m s-1 onto a rough aluminium surface Helmet site Max

linear acc. g

Max. FN kN

Max. FT kN

Max. z axis rot. acc.

krad s-2

Max. y axis rot. acc.

krad s-2

Max. liner crush mm

Loading slope

N mm-1 Avanti crown 148 6.5 1.1 3.5 < 1 13 oscill Avanti F 60° 125 5.5 1.0 -4.0 -1.8 16 330 Arc F 60° 129 6.2 1.0 -3.7 -3.5 14 390 Arc R 70° 138 6.5 0.5 3.2 2.1 * 14 480 Avanti R 70° 135 6.3 0.3 5.5 1.5 * 13 510

* x axis rotational acceleration

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 1 2 3 4 5 6

tang

entia

l for

ce F

T kN

normal force FN kN

Fig. 10. Variation of tangential with normal force at the shell/road interface for oblique impact of Arc helmet on the frontal 90° site.

0

1

2

3

4

5

6

7

0 5 10 15 20

head

form

forc

e

kN

liner deformation mm Fig. 11. Headform impact force vs. liner deformation for right 70° oblique impact of S1 helmet with VN = 5.4 m s-1 and VT = 3.6 m s-1 on flat surface. Deformation: o from minimum head to road distance, …… calculated by integrating headform acceleration.

the helmet microshell, implying that the shell then rolls on the rough surface. The final rotation of the helmet on the headform can be large about an ear-to-ear axis (Fig. 12a) but smaller about a neck to crown axis (Fig. 12b). In spite of differences in retention strap fixing locations (Table 2) there was no indication that the better systems reduced the overall helmet rotation in the impact, compared with the others. Therefore differences in retention systems seem only to influence helmet stability in non-crash situations.

6. Discussion

The test rig could simulate a wide range of oblique impact test parameters, and measure the headform linear and rotational accelerations. It was possible to simulate realistic helmet rotation by mounting helmets on a headform with a wig and a plasticized PVC scalp. In spite of the chinstrap being tight, and the lower jaw non-deformable, there was significant helmet rotation on the headform. The normal and tangential forces on the road surface were measured, for use [5] in determining the road/shell interface conditions. Therefore the first two aims of the introduction were achieved. In future, a more powerful actuator could provide a larger horizontal component of head velocity. The FEA [5] predicted headform rotational acceleration about all three axes, so the test rig should be fitted with a third rotational accelerometer. It is suggested that an oblique impact test, using a headform with a scalp and wig with measurements of rotational acceleration, is included in EN 1078. Variations in the external shape of helmets had only a small effect on the peak headform peak rotational accelerations. However, direct impact tests caused almost as large rotational accelerations as oblique impact tests with VH of 4 ms-1, for reasons explained in [5]. Although helmet liner densities have increased over the last 20 years to compensate for the presence of large ventilation slots, the peak headform linear accelerations have not been compromised; helmets provide greater protection since the impact energy in EN 1078 is 50% larger than that in BS 6863. The change in the peak headform rotational acceleration is unknown, since measurements were not made on early models of helmets.


a

b

Fig.12. Aventicum helmet positions after oblique impact tests: a) forwards rotation, when impacted on the front, b) lateral rotation, when impacted sideways on the crown.

The headform rotational acceleration was rarely

greater than 5 krad s-2, so it is unlikely that any diffuse brain injury would occur, if the criteria of rotational accelerations > 10 krad s-2 and rotational velocities > 100 rad s-1 [25] are valid. The impact site and direction affected the peak headform rotational acceleration; a rearwards impact on the side of the helmet caused higher values than a downwards impact on the helmet front. It is difficult to prove that helmets attenuate rotational head accelerations without carrying out comparable tests without a helmet. However the peak headform rotational accelerations when a helmet is worn, in tests representative of many bicycle crashes, are too low to cause brain injuries. Hence the criticisms of Curnow [13] and Franklin [14] are invalid. The peak linear acceleration measured using a headform with a relatively biofidelic scalp appeared similar to those typical when using a metal headform with no scalp layer in EN 1078. However, it would be necessary to test helmets containing instrumented cadaver heads to check Henderson’s criticism [15], and check the levels of peak linear acceleration for a 1.5 m drop height, compared with the use of a 5 kg metal headform.

It is likely that shear of the soft helmet interior padding, slip in the wig, and/or shear of the scalp lead to the low friction coefficient of circa 0.2 [5] at the helmet/headform interface. Although high frictional forces were measured [3] when no-shell helmets made oblique impacts on rough concrete surfaces, this is not the case for European helmets in 2006 impacting a roughened metal surface. Tarmac road surfaces have a smaller scale of roughness than concrete, while the local indentation of the exterior of a helmet with a micro-shell over a high-density EPS liner is relatively small.

Although the normal velocity component was not varied in the experiments, the oblique impact data confirms the loading slope design method for bicycle helmets [16], discussed further in [5]. This predicts the maximum impact force and liner crush increase in proportion to the normal velocity component, until the liner bottoms out. As the liners were about 75% crushed at the Right/Left 70° site, but only about 50% crushed at the crown impact site, the normal velocity component could be higher without the peak headform acceleration reaching the EN 1078 limit of 250 g. The same helmet must pass impacts on a kerbstone anvil at a 4.5 m s-1 velocity, and after conditioning at either -20 °C or at 50 °C. Consequently, to pass all the tests, there will be a considerable safety margin for the flat anvil tests on some sites at 20 °C. It would be advantageous to have a near-uniform liner thickness, rather than lower thickness at the sides (Table 1).


7. Conclusions

Oblique impact tests were performed that are representative of bicycle crashes. The headform instrumentation assessed the peak linear and rotational head accelerations, while the measured forces on the ‘road’ were useful for development of FEA models, allowing a better understanding of slip at the road-shell and helmet/head interfaces. Current helmet designs provide adequate protection for typical oblique impacts on to a road surface, in terms of the peak linear and rotational head accelerations. Most criticisms of current bicycle helmet designs are not valid: although test

headforms lack a deformable scalp, so have a high contact stiffness, this does not lead to inappropriate designs; there is a linear increase in the peak impact force with impact velocity, not a just sub-lethal level for minor impacts.

Acknowledgments

The authors thank EPSRC for support under grant R89790, and to the EPSRC Engineering Instrumentation Pool for the loan of a high speed camera.

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