N J Mills

S Wilkes

FEA of oblique impact tests on a motorcycle helmet

Keywords: motorcycle helmet, oblique impact, Finite Element Analysis

In accidents, motorcycle riders full-face helmets often make oblique impacts with road surfaces. Finite Element Analysis was used to predict the rotational and linear acceleration of a Hybrid II headform, representing a motorcyclist's head, in such impacts, considering the effects of friction at the head/helmet and helmet/road interfaces. Simulations of the oblique impact test in British Standard BS 6658 were validated by comparison with published data. This showed that COST 327 experimental data was largely determined by the friction coefficient (0.55) between the helmet shell and abrasive paper, and hardly affected by that between the head and helmet. Slip was predicted at the shell/paper interface throughout the impact, due to the high angular inertia of the helmet, and the normal force remaining below 3.5 kN. Simulations of more severe motorcycle helmet impacts explored the effects of impact site and direction, impact velocity components, helmet fit and the scalp. In these impacts, the higher velocity component normal to the road caused high frictional forces on the helmet shell, eventually causing it to roll on the road. The peak headform rotational accelerations, at some impact sites, were potentially injurious. The most effective method of reducing head rotational acceleration could be a reduction in the linear acceleration limit of the helmet standards.

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INTRODUCTION

Motorcycle helmets nearly always impact the road surface obliquely [START_REF] Otte | Technical demands on safety in the design of crash helmets: biomechanical analysis of real accident situations[END_REF], rather than perpendicularly, causing both linear and rotational acceleration of the head (these terms refer, throughout the paper, to the accelerations of the head's centre of mass). Excessive head rotational acceleration causes brain damage to animals [START_REF] Gennarelli | Diffuse axonal injury and traumatic coma in the primate[END_REF]. Research [START_REF] Depreitere | Mechanics of acute subdural haematomas resulting from bridging vein rupture[END_REF][START_REF] Yoganandan | Biomechanical aspects of blunt and penetrating head injuries[END_REF] suggests that the rotational acceleration must exceed circa 10 krad s -2 in the mid-sagittal plane (dividing

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2 symmetrical left and right halves) to cause subdural haematomas in humans. However, in motorcycle helmet test standards [START_REF]Protective helmets for vehicle users[END_REF][START_REF]of protective helmets and their visors for drivers and passengers of motorcycles and mopeds[END_REF], the majority of test impacts are perpendicular to the surface of fixed anvils, and the peak headform linear acceleration is used as a pass/fail criterion. This implies that the acceleration is unaffected by the head velocity component tangential to the anvil surface, and that peak linear acceleration determines head injury severity. A review of motorcycle helmet effectiveness [START_REF] Liu | Helmets for preventing injury in motorcycle riders[END_REF] showed that wearing a helmet reduced the risk of head injury by 72%, implying that helmet design was effective but not ideal. Attempts have been made to associate particular motorcyclists' head injuries with head translation and head rotation [START_REF] Richter | Head injury mechanisms in helmet-protected motorcyclists[END_REF], but headform rotational acceleration is not measured in helmet standards.

A review (chapter 16 of [START_REF] Mills | Polymer Foams Handbook-Engineering and biomechanics applications and design guide[END_REF]) concluded that Finite Element Analysis (FEA) of direct impacts, on helmets containing a near-uniform thickness of expanded polystyrene bead foam (EPS) liner, was reasonably successful. However, for at least 10 years, nearly all helmets have had softer structures in the crown region (holes or grooves in a single density liner, or a lower density EPS or other foam as a crown insert), a feature not considered in the FEA. Helmet shells are stiffest when loaded at the crown [START_REF] Gilchrist | Impact deformation of ABS and GRP helmet shells[END_REF], since that site has a doubly-convex curvature and is distant from any free edges. The softer liner in the crown region compensates for the high shell stiffness, and attempts to make the helmet impact response independent of site. The earlier FEA assumed that the EPS liner was bonded to the shell, whereas in practise it is either unbonded or only bonded in a small area at the crown. Glass fibre reinforced polyester resin (GRP) is used for many helmet shells, but its elastic properties depend on the processing conditions. Researchers often used GRP Young's moduli measured using flat test coupons, rather than values deduced from tests on helmet shells; Kostoupoulos et al. [START_REF] Kostopoulos | FEA of impact damage response of composite motorcycle helmets[END_REF] used an inplane Young's modulus of 19.7 GPa for a 2 mm thick prototype GRP shell. Aiello et al [START_REF] Aiello | Numerical simulations of motorcycle helmet impact tests[END_REF], who simulated direct impacts on a full-face Dainese helmet with a mixed-fibre composite shell, used an in-plane Young's modulus of 24 MPa at 50 °C. The consequent overestimation of the shell bending stiffness increased its load spreading to the EPS liner (the volume undergoing a high compressive strain), and hence gave erroneous predictions of helmet performance. The FEA often simplified the helmets, omitting chin-straps and foam inside the chin bar (which affect helmet rotation on the head), and the headform scalp.

No FEA has been reported of motorcycle helmet oblique impacts. In this multi-body problem, slip or rolling is possible at both the helmet inner and outer surfaces. The non-spherical head does not exactly fit the helmet liner, and the helmet liner does not exactly fit the helmet shell.

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Prototype helmets [START_REF] Halldin | A new oblique impact test for motorcycle helmets[END_REF], [START_REF] Phillips | European Patent EP1404189 Protective headgear and protective armour and a method of modifying protective headgear and protective armour[END_REF] aim respectively to reduce the friction inside and outside the helmet shell, by introducing easy-shear layers. The developers argued that this would reduce head rotational accelerations. However, such claims need to be demonstrated either experimentally, or by FEA.

There are few reports of headform acceleration measurements during motorcycle helmet oblique impacts; Aldman et al. [START_REF] Aldman | Oblique impacts -a parametric study in crash helmets[END_REF] dropped a complete, helmet-wearing dummy onto a 1 metre diameter rotating road surface. The vertical velocity was 5.2 ms -1 (typical of free fall from a riding head height of about 1.5 metres) while the horizontal velocity component was 8.3 ms -1 . The peak headform angular accelerations ranged from 7 to 14 krad s -2 , with higher values on an abrasive paper surface than on a surface of rounded stones. The peak headform linear accelerations ranged from 105 to 180 g, the same as for vertical drops at the same velocity onto a road surface.

The oblique impact test rig of BS 6658 [START_REF]Protective helmets for vehicle users[END_REF] uses a flat steel anvil, covered with a sheet of close-coated alumina abrasive paper, inclined at 15° to the vertical. The headform and helmet have a vertical velocity of 10 ms -1 at impact; the velocity components normal and tangential to the anvil surface are respectively V N = 2.59 ms -1 and V T = 9.66 ms -1 . Although V T may be representative of some motorcycle crashes, V N is much smaller than for typical falls. The test was introduced in 1985 to render obsolete meetings of expert assessors, who determined whether helmet geometries met the requirements of the previous standard BS 2495 [START_REF]Protective helmets for vehicle users (high protection)[END_REF]. It was not intended to simulate typical crashes, and headform accelerations were not measured.

The COST 327 project [START_REF]COST 327[END_REF] used this test rig for oblique impacts on the sides of four types of motorcycle helmets with thermoplastic or GRP shells. The Hybrid II headform used had a plasticized PVC 'scalp' but no hair. Its 570 mm circumference was just smaller than the nominal 580 mm circumference of the helmets tested, so it was a good fit to the helmet liners.

As the tests were limited to one site and direction, the range of peak headform rotational accelerations (2.4 to 8.5 krad s -2 ), for impact velocities from 6 to 12 m s -1 , will be less than that in real crashes. The friction coefficients for the helmet/abrasive paper interface ranged from 0.4 to 0.6.

FEA of bicycle helmet oblique impacts on a rough metal surface representing a road [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF] revealed that the effective friction coefficient of the helmet was a function of the friction coefficients λ R at the road/shell and λ H at the head/liner interfaces. The results of oblique impact experiments [START_REF] Mills | Oblique impact testing of bicycle helmets[END_REF], using an Ogle headform fitted with an acrylic wig, were replicated using λ R = 0.25 and λ H = 0.2. Bicycle helmets typically have much lower masses (0.3 kg) and

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angular moments of inertia (9 to 13 kg cm 2 ) than test headforms (typically 4 to 5 kg and circa 200 kg cm 2 ). Typical full-face motorcycle helmets have roughly twice the mass and twice the angular inertia of open-face helmets; the masses ranged from 0.8 to 1.7 kg, and the angular moments of inertia from 88 to 250 kg cm 2 [START_REF] Mills | The biomechanics of motorcycle helmet retention[END_REF]. Their tangential impact velocity component in crashes is often much higher than those for bicycle helmets. Consequently the oblique impact responses of motorcycle and bicycle helmets are expected to differ.

One aim of the research was to estimate the peak head rotational accelerations when motorcyclists fall to the road. Comparison with estimated human tolerance levels [START_REF] Depreitere | Mechanics of acute subdural haematomas resulting from bridging vein rupture[END_REF][START_REF] Yoganandan | Biomechanical aspects of blunt and penetrating head injuries[END_REF] and injury statistics should improve the understanding of head injury mechanisms. A second aim was establish the headform features and test conditions needed to realistically simulate crashes, hence to review the oblique impact tests in helmet standards. A third aim was to identify the factors that control the peak headform rotational acceleration in oblique impacts.

The vertical velocity component largely determines the amount of EPS liner crushing, hence the peak headform linear acceleration [START_REF] Mills | Polymer Foams Handbook-Engineering and biomechanics applications and design guide[END_REF]. Some sliding is expected to occur at the helmet shell-road and helmet-hair interfaces, so the friction coefficients should affect the peak tangential force on the head, which contributes to the head rotational acceleration. For bicycle helmets [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF], the headform geometry relative to the helmet liner geometry (which changed as the foam crushed) affected the normal force distribution on the interface, hence influenced the peak headform rotational acceleration. This mechanism may also apply for motorcycle helmets.

FEA Helmet geometry

A full-face motorcycle helmet with a GRP shell, that met EC Regulation 22 [START_REF]of protective helmets and their visors for drivers and passengers of motorcycles and mopeds[END_REF], was manufactured in 2001 by Mavet SRL, Campodoro, Italy. Its brand name was 'Dainese', its size 58 cm and its total mass 1395 g. The AGV full-face helmets (mass 1395 g, GRP shell) used in the COST 327 tests were no longer available, but a similar one was purchased on eBay in a nearly new condition; its geometry and foam densities were very similar to the Mavet helmet. Therefore the comparison of the FEA predictions with the COST test results is justified.

A medical X-ray computer tomography (CT) scanner at EMPA Dübendorf was used on the helmet after the chin-bar foam components were removed. The pixel size was 0.40 mm in the horizontal plane with 0.60 mm between slices. Such scans contain artefacts, especially from

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5 metal components such as the steel 'hanger plates' to which the chin-straps are attached. A horizontal slice image (Figure 1) shows small gaps between the EPS liner and the GRP shell near the hanger plates. The liner is an interference fit in the shell, and is not bonded to it. The relative positions of the helmet components are visible in the CT scan. However, the image quality was inadequate for satisfactory extraction of the geometry for the two densities of EPS liner; their grey levels varied spatially (possibly due to the air channels between the beads) and overlapped with the grey levels of other components.

AMIRA software [START_REF]AMIRA 4 software[END_REF] was used in Birmingham University Archaeology Dept. to improve the image contrast. The extracted inner and outer surfaces of the helmet shell contained many small holes, artefacts of the image analysis, and included the steel hanger plate surfaces.

Therefore Magics 12 software [START_REF]MAGICS 12 software[END_REF] was used to repair and simplify the shell geometry. Since the helmet shell was modelled in FEA by shell elements rather than by 3-D solid elements, only the shell inner surface was preserved. Much of the detail of the GRP weave on this surface was removed by localised smoothing and reducing the number of triangles in the .stl file to 6900. Figure 2 shows the helmet shell, as part of the complete FEA model. Small ventilation holes in the helmet liner were filled (forehead ventilation uses complexshaped holes, made by bonding an EPS insert to the main liner) before it and the chin bar foam components were laser scanned in Birmingham. Subsequently, the liner was cut vertically at one side with a band-saw, and the low density EPS crown moulding (which had been moulded through a hole in the crown of the main liner) was separated. The main part of the internal surface, where the two densities of EPS met, was part of an ellipsoid, with, at its top, radius of curvature 120 mm in the xy plane of Figure 3, and 100 mm in the yz plane. The vertical sides of the internal surface were part of a cylinder of elliptical cross-section, with half axes 94 and 78 mm. These surfaces were constructed in a CAD program Rhino [START_REF]Rhinoceros 3 software[END_REF], then used in ABAQUS [24] to separate the liner into two regions, which could be allocated different properties.

An initial chin-strap shape was created in ABAQUS CAE. A 25 mm wide strip, of radius of curvature 53 mm, was a close fit under the headform chin. It was linked on each side with two planar segments that passed around the cheek mouldings, and through the holes in the hanger plates. A preliminary ABAQUS run pulled the strap ends up through the hanger plates, until the strap conformed to the face and chin bar side mouldings. The shape of this deformed mesh was then imported into the main ABAQUS model; when used in the main model, it was initially stress free.

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Helmet components and their masses

Figure 3 shows the helmet foam components on the headform. The hanger plate positions, and chin-strap passage around expanded polypropylene bead foam (EPP) cheek mouldings, are typical of helmets that pass the 'retention (detaching) test' of Regulation 22-05. The shell mass was 615 g after removing the chin-straps (50 g), hanger plates (26 g), plastic ventilation mouldings (23 g), visor mounts (21g), base and vision-opening trim (84 g). In Table 1 the 'material mass' is the product of the component volume and its density, while the 'nonstructural mass' accounts for items rigidly attached to the component, such as adhesives, plastic clips joining the chin bar foams, and shell trim. Since the shell thickness modelled was 1.50 mm (see next section), 69 g of non-structural mass was added to correct the shell mass, while a further 107 g is added for trim and ventilation mouldings. The helmet would be tested without 105 g of visor and visor fixings, so it would have mass 1289 g. 208 g of flexible open-cell polyurethane (PU) foam, cloth and flexible PVC are not modelled, since they do not react to helmet acceleration on the 10 ms time scale of the impact (in subsidiary modelling, the effect of including the PU foam cheek pads, with material properties from the literature [START_REF] Mills | Modelling the indentation of low density polymer foams[END_REF][START_REF] Mills | Finite Element models for the viscoelasticity of open-cell Polyurethane foam[END_REF], was a less than 1% change in the peak accelerations). For the same reason, the circa 3 mm thick layer, of cloth backed with PU foam, that fits inside the liner was not modelled.

FEA showed that the model helmet had angular inertias 145, 145 and 135 kg cm 2 about earto-ear, crown-to-neck and nose-to-rear axes respectively (1 kg cm 2 = 10 -4 kg m 2 ).

Helmet material properties

A form of reverse engineering was used to characterise the GRP helmet shell modulus.

Sectioning the shell revealed a gel coat plus paint layer circa 40 μm thick, which was assumed to be insignificant for modelling purposes. There is thicker gel coat (polyester thermoset) near the sharp internal corners of the vision opening, which has Young's modulus circa 3 GPa.

The shell thickness was remarkably uniform, a result of pressure bag moulding; it varied from 1.3 to 1.5 mm over all but the chin bar and visor mount regions, but was 2.5 mm in the centre of the chin bar and circa 2 mm near the visor mounts. The shell density in the crown and chin bar areas was respectively 1830 ±10 and 1670 kg m -3 . For a GRP shell of uniform density 1830 kg m -3 and area 0.1990 m 2 to have a mass of 615 g, its mean thickness would be 1.69 mm.

Experimental load-deflection relationships were measured quasi-statically for the three loading geometries shown in Figure 4. FEA was then performed for the same loading geometries; a GRP Young's modulus E G = 8 GPa and Poisson's ratio 0.1 [START_REF] Lin | Failure analysis of woven-roving composites with moulded in holes[END_REF], with shell

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7 thickness 1.50 mm, reproduced the experimental responses (Figure 5) within 5%. The peak force in the experimental crown-loading data occurred when ring shaped cracks appeared in the resin; data beyond this point should be ignored. There was a slightly larger error for crown loading; the stiffness there depends critically on the highly loaded area near the visor mounts.

GRP of density 1830 kg m -3 would have an unrealistically high glass volume fraction V f = 0.48, if the resin density was 1200 kg m -3 (the glass fibre density is 2540 kg m -3 ). The resin is probably filled with a mineral powder to reduce shrinkage, hence its density exceeds 1200 kg m -3 . A more realistic V f = 0.3, with E Glass = 80 GPa and E resin = 3 GPa, leads to a fibredirection Young's modulus in unidirectionally-reinforced GRP of E u = 26.1 GPa. The woven roving layers in the shell produced a material with near-isotropic in-plane Young's modulus of 8 GPa that is close to the expected [START_REF] Hull | Introduction to composite materials[END_REF] 3/8 E u , i.e. 9.8 GPa. Modelling of crown impacts of GRP shell helmets onto a 50 mm radius rigid hemisphere at 7.5 ms -1 [START_REF] Kostopoulos | FEA of impact damage response of composite motorcycle helmets[END_REF] predicted a small region of delamination in the GRP; nevertheless 49% of the energy was stored elastically in the GRP compared with 8% dissipated by delamination. When GRP shell helmets are examined after BS 6658 test impacts on to flat anvils, no evidence of delamination is found.

Therefore, as only flat surface impacts are considered here, failure mechanisms in the GRP were not modelled. The 1.3 mm thick polyethyleneterephthalate PET fibre webbing chin-strap had material constants given in Table 2. These were obtained [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF] by approximating the experimental tensile response of a length of chin-strap taken from a helmet.

Sections were cut from the EPS and EPP foam components with a band saw, removing the moulded surface, which is usually of a higher density. Their densities (Table 2) were measured with an electronic densimeter, using the Archimedes principle. The uniaxialcompressive response of low density closed-cell polymer foams on loading can be fitted with

R P C C - - + = ε ε σ σ 1 0 0 (1)
where σ C and ε are the engineering compressive stress and strain, P 0 the effective gas pressure in the cells, and R the foam relative density (the foam density divided by the density of solid polymer). Impact compression stress -strain parameters for the EPP [START_REF] Mills | Time dependence of the compressive response of polypropylene bead foam[END_REF] are given in Table 2.

Stauffer [START_REF] Stauffer | Sandwich-Strukturen für ein besseres Dämpfungsverhalten in Polystyrolbasierten Helmen[END_REF] showed that the impact compressive stress, of EPS of densities 15, 30 and 50 kg m -3 , was 20% higher for an impact velocity of 9 ms -1 , than it was in a compressive test at 75 mm s -1 . Therefore compression tests at a crosshead speed of 20 mm min -1 were performed on cubes of side approximately 25 mm cut from the EPS foams, and the data fitted with equation 6 shows that this provides a good fit to the post-yield loading data. The σ C0 values from the straight line fit were increased by 20% before entering into Table 2. These values are consistent with literature impact data for the respective densities [START_REF] Hazarika | Stress strain modelling for EPS geofoam for large strain applications[END_REF], [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF] of EPS. The foam Poisson's ratios in Table 2 are values for the pre-yield response, measured at low strain rates with the equipment described in [START_REF] Mills | Polymer Foams Handbook-Engineering and biomechanics applications and design guide[END_REF] chapter 5.

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The limitations of the crushable foam material model in ABAQUS are discussed in chapter 6 of [START_REF] Mills | Polymer Foams Handbook-Engineering and biomechanics applications and design guide[END_REF]; there is no hardening in simple shear, and the unloading after uniaxial compression is too sudden. It requires an input of the uniaxial compressive data, plus the ratios p t / p C0 of respectively the initial yield pressures in hydrostatic tension and compression, and σ C0 / p C0 . p t / p C0 = 1.0 and σ C0 / p C0 = 1.933 were used. The use of the model had been validated by comparing predictions with experimental data for compressive impact on truncated EPS pyramids [START_REF] Gilchrist | Impact deformation of rigid polymeric foams: experiments and FEA modelling[END_REF], and oblique impacts on bicycle helmets [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF]. Hence the limitations of the model do not cause significant errors for the type of deformation fields considered here.

Headform

A Hybrid II headform of width 160 mm was used at EMPA for helmet testing. The scanned shape of a 580 mm circumference headform had been used for bicycle helmet FEA [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF]. Its 199 mm length equalled that of the 70 th percentile adult male, but its 154 mm width was smaller than the 160 mm of 70 % adult male. Therefore two versions of the headform were used; the original 'narrow' version and a 'broad' version with width stretched to 160 mm using Rhino CAD software.

The test headform had an approximately 10 mm thick plasticized PVC scalp stimulant outside the aluminium casting. The response of a similar headform scalp was measured [START_REF] Mills | Oblique impact testing of bicycle helmets[END_REF] under plane strain compression conditions. The compressive stress vs. deflection graph up to a stress of 11 MPa was approximated by linear segments (Table 3) and used as the normal contact stiffness function of the head/helmet interface. This obviated the need to create and mesh the scalp geometry. The PVC scalp layer shear stiffness, measured in a slow shear test as 2.3 x 10 8 Pa m -1 , was used as the interface 'elastic slip stiffness'. In [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF] FEA predicted, for 200 mm drops onto a flat rigid table, peak linear headform accelerations at two sites that were within the range expected for biofidelic headforms.

The headform, with axes shown in Figure 3, had length 199 mm, breadth 154 mm, and the rotational moments of inertia of a test headform I xx 199 kg cm 2 , I yy 237 kg cm 2 and I zz 172 kg cm 2 [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF]. In a wire frame projection onto the mid-sagittal plane at the start of a simulation (Figure 7), a circa 3 mm gap is seen between the liner interior and the headform, which is

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filled in the real helmet with soft polyurethane foam. Hence there was no initial contact between the headform and the helmet liner. As there was no tension on the chin-strap at the start of the simulation, there was no initial compressive loading of the liner.

Contact conditions

Penalty contact algorithms were used for the interfaces; the default stiffness, normal to the interface, is ten times the underlying element stiffness.

At interfaces between the headform and the liner, cheek pads and chin bar foams, elastic slip (shear of the interface before slip occurs) simulated the shear of the scalp layer. The parameters for this and the normal stiffness function (Table 3) are given in the previous section. At the road/shell interface, a bi-linear contact stiffness had zero normal pressure at initial contact, 10 MPa at a 1 mm over-closure and 100 MPa at 2 mm over-closure. This reduced force oscillations to an acceptable level, while the < 1 mm peak over-closure was insignificant. Hard contact, with a friction coefficient of 0.5, prevented the liner outer surface sliding out of the shell inner surface. Elragi [33] reviewed conference papers and theses, and found 0.5 that the dynamic friction coefficient of low density EPS on itself exceeded 0.5.

There is no published data for higher density EPS sliding on other substrates. The friction coefficient at the helmet shell/anvil interface was taken from the analysis of the COST 327 tests, in which the initial contact force was small and sliding was continuous; hence it is a dynamic friction coefficient.

Meshing

In preliminary FEA, the part meshes were seeded at 15, 10, 7 and 5 mm spacing; the final change caused less than 1% change in the predicted responses, so a 5 mm spacing was used subsequently. Verification of the liner mesh showed that the average shape factor (element volume/ optimal element volume) was 0.62 and the average aspect ratio was 1.8. 4-node linear tetrahedral elements C3D4 were used for the foams; the helmet liner had 251571, the central chin bar foam 19928, and the cheek pads 16632 elements. The shell had 21317 linear triangular shell elements S3R, the chin-strap 803 linear triangular membrane elements M3D3, while the headform and hanger plates had respectively 4222 and 235 linear rigid triangular elements R3D3. ABAQUS indicated the critical elements, which determine the time interval, were in the liner; the region around them was re-meshed, after faces were merged using virtual topology, to ensure the time interval exceeded 150 ns. This allows a reasonable CPU time of about 5 hours.

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Impacts

The test rig coordinate system 2 axis (Figure 2) is normal to the anvil (road) surface, as is the impact velocity component V N . The tangential velocity component V T of the impact is along the 3 axis. The head orientations relative to the anvil are determined by a sequence of rotations, from an initial position with the crown touching the anvil and the face along the 3 axis (Table 4). The effect of the rotations 'Right 35 up at 45' are shown in Figure 2b. In the FEA headform acceleration components are determined in the 123 axes. Dynamic (explicit) ABAQUS was used for the 20 ms duration of impact, with the large deformation option.

COST 327 tests

The COST 327 motorcycle safety helmet project involved four European countries in accident surveys, the establishment of human injury tolerance, and the development of helmet test methods. Part of the last topic, carried out at EMPA, involved the development of oblique impact tests. A Hybrid II headform was equipped [START_REF]COST 327[END_REF] with nine accelerometers (Endevco The drop tests were performed onto an anvil inclined at 15° to the vertical, fitted to a steel block, fixed on a concrete block with a total weight of 1 tonne. The anvil, to the specifications of BS 6658, was equipped with a tri-axial Kistler type 9366AB force transducer, fixed on a mounting plate (230×300 mm), allowing the measurement of both normal and tangential force components. The grade 80 alumina coated paper, described in the introduction, was replaced after each impact. The accelerometer and force transducer signals, electronically filtered according to CFC600, were recorded for 25 ms. High speed video at 4500 frames/s was taken of one test. Data for an impact velocity of 10 m s -1 were used. The helmet (and headform) fall direction lay in its mid-sagittal plane, while its z axis (see Figure 3) was horizontal, and its x axis downwards.

FEA PREDICTIONS

Reproducing COST 327 tests

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The 'broad' headform of mass 4.77 kg was used for these simulations. The impact velocity components normal and tangential to the anvil were V N = 2.59 m s -1 and V T = 9.66 m s -1 . The experimental overall friction coefficient was used as the shell/road friction coefficient λ R , while the unknown head/liner friction coefficient λ H was varied (Table 5). The predicted peak normal F N and tangential F T reaction forces for the Mavet helmet were close to the respective experimental values, in spite of the helmets differing slightly in design and materials. However, as these predicted values were insensitive to λ H , they did not allow the friction coefficient determination. The overall friction coefficients (Table 5), the slopes of straightline fits to F T vs. F N data, are close to λ R , and the correlation coefficients r are high, for all but λ H = 0.2. The peak headform rotational acceleration increased almost in proportion with λ H ; consequently a λ H of 0.5 best replicated the COST 327 data. The tangential impulse, the time integral of F T , increases slightly with λ H , and the experimental impulses are consistent with λ H = 0.5.

Figure 8 compares FEA predictions of F T vs. F N with experimental data. Both plots were nearly linear, implying that the shell slid on the anvil for the whole impact; this was confirmed by comparing (Figure 9) frames from the experimental movie with the FEA simulation at four times. After shell-to-anvil contact, the headform side approached the anvil, and both headform and helmet rotated. The oval shape of the shell base distorted significantly during the impact. Slip was observed at the head/liner interface in the FEA simulations for λ H = 0.2, but not when λ H = 0.4.

FEA predicted the force F N had a small initial peak, then reduced somewhat as the shell partly rebounded from the anvil, before a main peak (Figure 10a). The magnitude and shape of the main peak agreed well with data for the AGV helmet. However the first peak is smaller than, and the main peak 2 ms later than, the experimental data. This time delay suggests the initial separation between the side of the model headform and the liner is 5 mm greater than in the experiments. Increasing the helmet shell mass locally near the impact site would increase the size of the first peak. The coupling of the helmet and headform masses can be assessed using plots of the reaction force magnitude vs. the linear acceleration magnitude (Figure 11). The initial portions of these graphs are non-linear, because the helmet shell accelerated before the liner impacted the

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headform. The slope of the linear regression line can be used to evaluate the effective mass m e of the helmet-headform combination. The 6.19 kg value for the COST 327 experiments exceeded the 5.36 kg for the FEA model (the respective correlation coefficients were 0.901 and 0.938). Since the total headform and helmet mass was 6.06 kg, the helmet and headform masses were well coupled in the experiments, probably due to tension in the chin-strap. In the FEA model no initial tension was applied to the chin-strap.

Figure 12 shows contours of the equivalent plastic true strain PEEQ on both surfaces of the liner, when the plastic zone size was maximal after 11 ms. The yielded zone is quite small and nearly circular on both surfaces. On the inner surface there is some yielding of the low density EPS insert, and the zone is limited by the rear ventilation groove.

In a λ H = 0.5 simulation of a 257 g lighter helmet (without the non-structural mass of Table 1, so with moments of inertia 105, 109 and 100 kg cm 2 ), the ratio F T / F N reduced after 15 ms, showing that the shell rolled on the anvil at the end of the contact. Therefore the helmet angular inertia must exceed a critical value to maintain slip at the shell/road interface. This critical value depends on the impact site, and the friction coefficients. In higher V N simulations, which caused higher peak forces F N , rolling was common and the plots were non-linear. Therefore, shell sliding only persisted on the abrasive anvil under COST 327 impact conditions because F N remained relatively low and the full-face helmet angular inertia was high.

The correlation coefficients r between the rotational acceleration magnitude θ &

& and the tangential force F T were 0.86 and 0.88, for the AGV helmet data and the FEA simulation respectively, because the time dependences of the two variables differed. As with bicycle helmets [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF], the line of action of the net force F N from the large contact area does not pass through the headform centre of gravity, so F N contributes to 1 θ & & . However, as the neck-to- crown axis rotational inertia of the helmet (143 kg cm 2 ) was of similar magnitude to that of the headform (163 kg cm 2 ), and the helmet could slip on the headform, the headform rotational acceleration could not be simply linked to the sum of these two contributions.

Simulation of typical oblique impacts in crashes

The validated FEA model was used to explore more severe impacts, simulating serious motorcycle crashes. A headform/liner friction coefficient of λ H = 0.2, and a shell/ road friction coefficient λ R = 0.4, were found in bicycle helmet oblique impacts [START_REF]COST 327[END_REF][START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF] using a wig on the

T P I R C S U N A M D E T P E C C A ARTICLE IN PRESS
13 headform, and a roughened metal plate as the road surface. The headform mass of 4.26 kg represented the human head. The 'narrow' headform shape used is typical of an imperfect fit of motorcyclists' heads to their helmets [START_REF] Gilchrist | Survey of head, helmet and headform sizes related to motorcycle helmet design[END_REF]; some have helmets that are tight at the sides and others at the front/rear. In the simulations, either one side of the chin-strap, or both sides for crown impacts, initially slackened as the EPS liner crushed. Later, for the frontal impacts, the chin-strap tightened after the helmet had rotated on the head, but the total chin-strap force was never high.

First, the impact site was varied, with a velocity V N of 5 m s -1 representing a fall from a riding head height of about 1.5 m. The sites were chosen to induce rotations about the three headform axes. As the rider's shoulder probably hits the road first for impacts below left/right 70° sites, a right 80° site was chosen, in place of the COST 327 impact site. The peak rotational headform accelerations, given in Tables 6 and7, were smallest for the frontal 80° site because the contributions from F T and F N acted in opposite directions, as for bicycle helmets [START_REF] Mills | Finite element analysis of bicycle helmet oblique impacts[END_REF]. They were highest for the right 80° site, where the impact direction caused rotation about the headform crown-to-neck axis.

The maximum values of F N and linear headform acceleration hardly changed as V T increased from 5 to 10 m s -1 . However, max 1 θ & & increased somewhat with V T , and exceeded the 10 krad s -2 level for the onset of diffuse axonal injury for the right 80° and the 'right 45° up at 45°' sites.

The value of F T /F N was affected by the impact site, and was only close to λ R when V N = 10 ms -1 .

The minimum thickness of the EPS liner (Tables 6 and7) was calculated from the closest distance between the headform and the road, allowing 1.5 mm for the shell thickness. At the crown and right 45° sites, this was more than half the original liner thickness, showing that the helmet had considerable protection in reserve. For impacts on the front of the helmet, the chin bar foam came into play; the predicted accelerations are for a headform with a rigid chin, rather than an articulated jaw.

Figure 13a and b compare the time dependence of the headform accelerations for the right 80° impact site and V N = 5 ms -1 , for three values of the tangential velocity V T . As V T increased, the linear acceleration peaks at 4.5 and 7.5 ms decreased in size slightly, while the acceleration was slightly greater at times > 10 ms. In contrast, the initial rotational acceleration peak increased significantly in size and duration as V T increased.

T P I R C S U N A M D E T P E C C A
Second, a range of normal and tangential velocity components was used for the right 80° site (Table 7). The peak liner crush distance increased approximately in proportion to the velocity V N . The peak linear headform acceleration is estimated to reach the 275 g limit when the minimum liner thickness is 15 mm; hence the liner density is higher than optimum for impacts onto flat surfaces. Even in direct impacts, there were significant rotational accelerations when V N ≥ 5 ms -1 . In the oblique impacts, the peak rotational acceleration increases significantly with V N , but hardly with V T , while the variation of peak rotational acceleration (krad s -2 ) with peak linear acceleration (g) could be described by

1098 2 . 98 max max - = a θ & & (2)
with correlation coefficient r = 0.986. This relationship was similar to, but for greater a max values than, that reported in [START_REF]COST 327[END_REF] for the COST 327 impact site, with coefficients 75.4 rad s -1 g - 1 and +500 rad s -2 , and r = 0.91.

In an individual high-velocity test (with parameters given by the 9 th row in Table 7), the correlation coefficient between the tangential force F T and θ & & was only r = 0.72. However, the correlation between the maximum values, for a series of tests on nearly the same impact site, was reported [START_REF]COST 327[END_REF] as r = 0.97, suggesting that the easily-measured F Tmax could be used as a surrogate for the more-difficult-to-measure max θ & & . Figure 14 shows, for the FEA simulations in Tables 6 and7, a good correlation for a single site (r = 0.945 for the right 80° site), but no overall pattern for a range of impact sites. For lateral impact sites, the reported [START_REF]COST 327[END_REF] 3.17 slope of the best-fit line for a number of helmets is half the slope of 6.34 in Figure 14 for a slightly different impact site. Therefore measurements of F Tmax are not surrogates for measurements of headform rotational acceleration.

The reported correlation [START_REF]COST 327[END_REF] between the tangential impulse J T (the time integral of F T ) and the peak headform rotational velocity was r = 0.95. When J T values (Ns) from Tables 6 and7 were plotted against V T (m s -1 ); the best-fit relationship was found to be

J T = 1.52 V T -0.57 (3)
with r = 0.89. The basis of this correlation is discussed later.

A shell/ road friction coefficient λ R = 0.2 was used to simulate helmets with low-friction surface layers. However, this reduction in λ R caused minor increases in the peak rotational accelerations at low V N , and minor increases at high V N (penultimate column in Table 7). For

V N ≥ 5 ms -1 , the main F N peak, at about 5 ms into the impact, caused both large flat areas to form on the outsides of the shell and liner, and the liner inside to conform to the headform

T P I R C S U N A M D E T P E C C A ARTICLE IN PRESS 15
surface over a similar area. The peak rotational acceleration occurred at this time. Therefore large geometry changes, caused by the high impact force F N , influenced the peak rotational acceleration more than the value of λ R .

When simulations were repeated for the wider COST 327 headform, the results, in the last 3 rows of Table 7, were similar to those for the narrow headform. Therefore headform fit does not greatly influence the peak headform accelerations.

DISCUSSION

The FEA modelling results will be discussed prior to the three research aims. A comparison, of compression tests on a complete shell with FEA of the same loading geometry, showed that the Young's modulus of the GRP in the Mavet helmet was relatively low. This is almost certainly true for other helmets, suggesting that the moduli used in [START_REF] Kostopoulos | FEA of impact damage response of composite motorcycle helmets[END_REF] and [START_REF] Aiello | Numerical simulations of motorcycle helmet impact tests[END_REF] were probably too high. GRP fracture and delamination does not occur in the flat anvil impacts, so these phenomena were not modelled. However, to model impacts on kerbstone anvils, these phenomena should be included in the FEA, as in [START_REF] Aiello | Numerical simulations of motorcycle helmet impact tests[END_REF]. If FEA is to replace prototype testing, in developing helmet designs to pass EC Regulation 22 or BS 6658, it is vital to model all the main protective foam components; the cheek pad foam was loaded in impacts at the helmet side, and the chin bar centre foam was loaded in frontal impacts. However, it is reasonable to replace the thin layer of soft polyurethane foam, which lies between the EPS liner and the headform, by an air gap. FEA predictions were close to the data (F N vs. time, F T vs. F N , and headform acceleration vs. time) from COST 327 oblique impact experiments [START_REF]COST 327[END_REF]. Some minor differences remain, due to the non-availability of the tested helmet, and doubts about how tightly the headform was strapped into the helmet. The test conditions caused sliding to continue on the abrasive paper surface throughout contact. The good fit of the test headform to the helmet liner, the high friction at this interface, and the high helmet moment of inertia, were all factors that prevented sliding at the headform/helmet interface.

The BS 6658 oblique impact conditions, chosen by Glaister [START_REF] Glaister | A test for the sliding resistance of protective helmets[END_REF] to test helmet shell protrusions within the limitations of a drop tower, are less severe than the majority of injurycausing motorcycle crashes. Consequently, the COST 327 program did not evaluate protection against excessive rotational head acceleration in more-typical crashes. For rearwards impacts on the right 80° site, FEA predicted that, for typical V N ≅ 5 ms -1 , 10 > V T > 5 ms -1 and 0.4 > λ R > 0.2, the peak headform rotational accelerations ranged from 14 to 15 krad s -2 . For some other helmet impact sites, the peak rotational accelerations for these

T P I R C S U N A M D E T P E C C A ARTICLE IN PRESS 16
velocities (Table 6) are less than the estimated 10 krad s -2 for the onset of brain injury. A survey of motorcyclist head injuries [START_REF] Richter | Head injury mechanisms in helmet-protected motorcyclists[END_REF] showed a quite high incidence of subdural haematomas, which can be caused by head rotation.

The headform used for rotational acceleration measurements in oblique impact tests should have a compliant scalp, and possibly a synthetic wig to simulate the hair, together with soft under-chin tissue. These features will lower the head/helmet friction coefficient compared with COST 327 experiments, hence allow helmet rotation. An impact velocity V N in excess of 5 ms -1 would ensure that the normal component of impact force was typical of serious helmet impacts. Oblique impacts, of an instrumented headform wearing a helmet, could measure the friction coefficients of road surfaces; this is likely to be less that that for the abrasive paper in BS6658.

The FEA showed that the shell/road friction coefficient was not the only factor determining the peak headform rotational acceleration in typical crashes. The deformation of the helmet liner, and the distribution of normal stresses across its inner surface (which depends on the impact site and direction), also make substantial contributions. The latter factor is related to helmet fit; real heads typically fit helmet liners less well than headforms, on which some liners appear to be based. The results in Table 7 suggest that a low shell/road friction coefficient, proposed for improved helmets [START_REF] Halldin | A new oblique impact test for motorcycle helmets[END_REF][START_REF] Phillips | European Patent EP1404189 Protective headgear and protective armour and a method of modifying protective headgear and protective armour[END_REF], had little effect on the peak head rotational accelerations for lateral impact sites. Finan et al [START_REF] Finan | The influence of reduced friction on head injury metrics in helmeted head impacts[END_REF] found reductions in the acceleration, for a 'Front 45 down' site in the current nomenclature, when a low friction layer was added to the outer surface of an American football helmet, but an increase in acceleration for the equivalent 'rear 45 up' site. Hence the benefit, if any, of a low friction layer is site specific. Rolling on the shell/road interface caused the tangential force F T to fall below that calculated from F N and a single friction coefficient, and the helmet geometry changed in the impact region. Given these factors, and the complex nature of the load spreading from the shell to the foam components, FEA is the only method to predict the rotational acceleration experienced by motorcyclists' heads in typical crashes.

The abrasive-anvil oblique impact test in BS 6658 requires that the tangential force does not exceed 4 kN, while its time integral does not exceed 28 Ns. When the same equipment is used in Regulation 22-05, the impact velocity is reduced to 8.5 ms -1 , and the limiting parameters reduced to 3.5 kN and 25 Ns respectively. However the fourth from bottom row of Table 7 shows that the headform rotational acceleration could be 20 krad s -2 , for a helmet test that met the tangential force and force integral requirements, and the Regulation 22-05 requirement for a peak linear acceleration < 275 g in direct impacts with V N = 7.5 ms -1 . A simple model 15) can be used to estimate the tangential impulse J T from the moments of inertia of the helmet and headform, and explain why J T correlates with V T . Assuming that no slip occurs at the head/helmet interface, and that the helmet radius r remains constant during the contact (it typically varied from 138 mm to 118 mm in a COST type impact), the angular equivalent of the impulse -momentum equation for linear motion is ( )

T P I R C S U N A M D E T P E C C A ARTICLE IN PRESS 17 (Figure
1 2 ω ω - = I J r T (4)
The contribution of F N to the headform angular velocity change is assumed to be zero. The initial angular velocity of the headform ω 1 = 0 while the final value ω 2 is assumed to be that for rolling on the anvil. Ignoring any reduction in V T during the impact (it typically drops from 9.66 to 6.9 ms -1 ),

ω 2 = V H /r so 2 r V I J T T ≅ (5)
Therefore, substituting the combined inertia of the Mavet helmet and headform I zz = 0.030 kg m 2 and r = 0.13 m, the tangential impulse was estimated as 18 Ns, 12% higher than the values in Table 5 for the COST data and FEA simulation. For the same input data, equation [START_REF]Protective helmets for vehicle users[END_REF] implies that the tangential impulse is proportional to 1.78 V T , which is 17% greater than the relationship in equation [START_REF] Depreitere | Mechanics of acute subdural haematomas resulting from bridging vein rupture[END_REF]. It seems unlikely that the 28 Ns limit, specified in BS6658, could be reached, because typical helmets have insufficient angular inertia. The purpose of the limit, in terms of head injury mechanisms, is unclear.

Consequently, a review of the standards may be in order. Equation [START_REF] Gennarelli | Diffuse axonal injury and traumatic coma in the primate[END_REF] showed the peak headform rotational acceleration for the right 80° site was almost linearly related to the peak linear acceleration. A reduction in the latter, from 275 g towards 150 g, should proportionately reduce the peak rotational head acceleration. This could be achieved, without changing the helmet shell dimensions, by reducing the liner foam density, if the impact velocity into a kerbstone anvil was reduced from the current 7.5 m s -1 . This kerbstone impact test critically influences helmet design, because it causes a greater liner crushing distance than flat anvil impacts, henceit determines the liner density. Since the most common crash type ison flat surfaces, flat anvil impacts should have the greatest influence on helmet design.

Consequently, a reduction in the kerbstone impact velocity should improve helmet protection for the majority of motorcyclists.

T P I R C S U N A M D E T P E C C A T P I R C S U N A M D E T P E C C A T P I R C S U N A M D E T P E C C A T P I R C S U N A M D E T P E C C A ARTICLE
P I R C S U N A M D E T P E C C A ARTICLE IN PRESS
T P I R C S U N A M D E T P E C C A ARTICLE IN PRESS
T P I R C S U N A M D E T P E C C A ARTICLE IN PRESS

  7264B-2000), positioned on a mounting block in a 3-2-2-2 array, following the recommendations of Padgaonkar et al. [34]. The total mass, including mounting block and the accelerometers, was 4.77 kg. The accelerometer signals were fed to three voltage amplifiers (Endevco Model 136) and sampled at 100 kHz using two Nicolet BE490XE transient recorder boards.

Figure

  Figure 10b and 10c compare respectively the magnitudes of the total linear and total rotational accelerations with COST tests on the left and right sides of two helmets; the agreement is good, allowing for a 2 ms delay in the predicted responses.

Fig. 1 .

 1 Fig. 1. CT scan slice of Mavet helmet at level of the upper rivet on the hanger plate.

Fig. 2 .

 2 Fig. 2. Complete Mavet helmet in the impact positions a) right 80, reached by a 80° rotation about the 3 axis from a crown impact site, b) right 45 up at 45. The tangential velocity component, along the 3 axis, is towards the viewer.

Fig. 3 .

 3 Fig. 3. FEA model of Mavet helmet with its shell removed, revealing the foam components and chin-strap position. The x y z axes of the headform are shown.

Fig. 4 .

 4 Fig. 4. Loading experiments for shell: crown (top), side, and chin.

Fig. 5 .

 5 Fig. 5. Comparison of experimental and FEA predictions for shell loading.

Fig. 6 .

 6 Fig.6. Compressive loading and unloading data for the two densities of EPS from the Mavet helmet, plotted according to equation[START_REF] Otte | Technical demands on safety in the design of crash helmets: biomechanical analysis of real accident situations[END_REF]. The dotted lines are fits of the post-yield, loading response, for the range 0.1 to 5 of the strain parameter.

Fig. 7 .

 7 Fig. 7. Wire-frame view of initial position of helmet on headform; gaps between the liner inside (grey line) and the headform are visible.

Fig. 8 .

 8 Fig. 8. F T versus F N for oblique impacts with COST 327 conditions: FEA with λ R = 0.55, λ H = 0.5, compared with data for AGV helmet.

Fig. 9 .

 9 Fig. 9. Helmet and headform positions in COST 327 oblique impact test with λ R = 0.55, λ H = 0.5, compared with experiment at times a) 0.6 ms, b) 6 ms, c) 12 ms, d) 18 ms.

Fig. 10 .

 10 Fig. 10. FEA predictions (solid curve) for λ R = 0.55, λ H = 0.5 of oblique impact of Mavet helmet, with COST 327 data from four experiments shown as dotted curves, for AGV helmet: a) normal force, b) linear acceleration, c) rotational acceleration magnitude.

Fig. 11 .Fig. 12 .

 1112 Fig. 11. Linear head acceleration vs. reaction force, comparing FEA with λ R = 0.55, λ H = 0.4 with COST data. The dotted lines are linear fits Fig. 12. Contours of true plastic strain PEEQ in the liner after 11 ms of the COST simulation with λ R = 0.55, λ H = 0.4.

Fig. 13 .Fig. 14 .Fig. 15 .

 131415 Fig. 13. Time dependence of a) peak linear acceleration and b) rotational acceleration, for three values of V T (ms -1 ), for a right 80° impact with V N = 5 ms -1 . Fig. 14. Peak rotational head accelerations vs. peak tangential forces for data in tables 6 and 7. Correlation line and dots for right 80 site, triangles front 80, squares crown, crosses r45 site. Fig. 15. Model for calculation of the tangential impulse.

T

  Figure 1

  Figure 7

Figure 15

 15 Figure 9

Table 1 .

 1 Helmet components in the FEA

	component	Material	Non-	Total	Associated soft	mass
		mass	structural	mass used in	component	(g)
		(g)	mass (g)	FEA (g)	ignored
	Main liner	204	40	244	Cloth liner	63
	Chin bar centre	13	29	42	Top, base cover 21
	Cheek pad x 2	12.5 x 2	6 x 2	18.5 x 2	Cloth, PU foam 42 x 2
	Chin strap	10		10	Covers, rings	40
	Hanger plate x 2 13 x 2		13 x 2
	GRP shell	546	176	722
	TOTAL	824	257	1081

Table 2 .

 2 Materials properties used in the FEA

	material	Density	Young's	Poisson's	Foam initial	effective
		(kg m -3 )	modulus	ratio	yield stress	gas pressure
			E (MPa)		 C (MPa)	P 0 (MPa)
	EPS crown insert	22	3.0	0.1	0.14	0.097
	EPS main & chin	59	19	0.1	0.62	0.19
	EPP cheek	70	20	0.1	0.76	0.27
	GRP	1830	8000	0.1
	PET webbing	700	1000	0.2

Table 3 .

 3 Normal contact response of the PVC scalp simulant

	pressure	Over-closure
	MPa	mm
	0	0
	0.63	1
	2.22	2
	4.6	3
	7.8	4
	11.2	4.8

Table 4 .

 4 Rotations of the inverted head, facing the 3 axis, to reach the impact site

	Site and direction	sequence of rotations
	Crown lateral	90 about 2 axis
	Front 80° down	80 about 1 axis
	Right 80° back	80 about 3 axis
	Right 45° up at 45°45 about 3 axis, 45 about 2 axis
	COST	180 about 2 axis, 90 about 3 axis, 15 about 1 axis

Table 5 .

 5 Oblique impacts on a flat abrasive anvil, using COST 327 conditions, with V N = 2.59 m s -1 and V T = 9.66 m s -1

	Head	Road	overall	Correl	max N F	max T F	Head	Head	impulse
	 H	 R		r	(kN)	(kN)	a max (g)	   (krad s -2 ) max	I T (Ns)
	AGV		0.549	0.997	3.66	2.06	77	5.5	16.2
	data.		0.016	0.001	0.22	 0.08	4	 0.4	0.7
	FEA 0.2	0.55 0.525	0.985	3.80	2.06	93	1.8	13.9
	FEA 0.4	0.55 0.549	0.9999 3.60	1.98	95	3.4	15.4
	FEA 0.5	0.55 0.549	1.0000 3.49	1.92	97	5.2	15.9
	FEA 0.7	0.55 0.550	0.9999 3.62	1.98	98	7.4	16.3

Table 6 .

 6 Oblique impacts on flat surface, using  R = 0.4,  H = 0.2, and V N = 5.0 m s -1 , for a range of impact sites and tangential velocities.

	Impact	V T	max N F	max T F	F T /F N (r) liner	head	head	head	head	head	impulse
	site & direction	(m s -1 )	(kN)	(kN)	x min (mm)	a max (g)	max 1    (krad s -2 ) (krad s -2 ) max 2    (krad s -2 ) max 3   	(krad s -2 ) max   	J T (Ns)
	Front 80°d	0	7.08 0.99	16.6 135 2.7	0.6	-0.9 2.7	1.7
	own	5	6.63 2.58 0.425 (0.972) 18.8 154 3.4	0.8	0.6	3.4	14.9
		10 6.75 2.68 0.400 (1.000) 18.1 154 3.7	0.9	1.3	3.8	17.8
	Crown	0	5.50 0.50	22.7 138 0.8	-1.0 2.0	2.2	-0.2
	lateral	5	5.10 1.72 0.199 (0.709) 22.4 134 5.2	-2.1 1.6	5.7	8.4
		10 5.09 2.04 0.366 (0.948) 22.4 143 6.2	2.9	1.9	7.0	14.2
	Right 45°U		7.79 0.49	25.4 163 3.8	-2.5 2.6	4.7	-2.2
	p at 45°0 5	8.04 2.04 0.217 (0.811) 26.7 170 10.7 -5.3 2.7	11.8	7.5

Table 7 .

 7 Oblique impacts, rearwards on the right 80° site, on a flat surface with  R = 0.4,  H = 0.2, for a range of velocity components 

	V N	V T	max N F	max T F	F T / F N (r)	liner	head	head	head	head	head	head	impulse
	(m s -1 )	(m s -1 )	(kN)	(kN)		x min (mm)	a max (g)	max 1    (krad s -2 )	max 2    (krad s -2 )	max 3    (krad s -2 )	(krad s -2 ) max   	 R = 0.2 max   	J T (Ns)
	2.5 0	2.62 0.30 -0.115 (0.50)	31.1 83	3.4	-0.6	3.1	3.9		-1.2
		5	2.54	0.94 0.360 (0.990) 31.5 68	4.3	1.4	2.9	4.8	5.3	6.3
		10 2.38 0.94 0.389 (0.999) 31.4 65	4.2	1.3	3.2	4.8	5.2	6.5
	5.0 0	8.03	0.40 -0.084 (0.63)	25.3 155 7.7	-1.0	6.8	10.3		-2.8
		5	7.12	2.24 0.252 (0.829) 25.6 145 13.2	2.9	6.3	14.3	13.6 7.7
		10 6.73 2.43 0.354 (0.992) 25.3 140 12.5	2.9	6.2	13.8	13.6 13.3
	7.5 0	11.95 0.53 -0.092 (0.65)	18.3 242 9.3	-1.6	8.1	12.2		-4.1
		5	11.10 2.77 0.14 (0.700)	18.7 226 18.1	3.0	8.3	19.5	19.0 8.2
		10 10.19 3.31 0.320(0.949)	18.5 208 18.4	3.7	8.0	19.9	19.8 18.5
	2.5* 10 2.74 1.04 0.381 (0.999) 29.3 59	4.8	1.1	2.9	5.1		8.2
	5.0* 10 7.15 2.39 0.338 (0.991) 22.6 131 11.9	2.3	5.7	13.0
	7.5* 10 10.11 3.06 0.310 (0.914) 14.7 205 16.6	3.6	7.0	18.1
	* using the wider headform of mass 4.77 kg

8.14 3.23 0.350 (0.949) 26.7 170 12.3 -5.8 -2.9 12.7 13.7

Acknowledgements

Dr M. Lobb of the Birmingham University Field Archaeology Unit is thanked for performing laser scans of helmet foam components. Dr P. Brüwihler of EMPA kindly provided the Mavet helmet, and commented on the paper.

Conclusions

The peak head rotational acceleration, when motorcyclists fall obliquely to the road, was estimated by FEA to be the order of 15 krad s -2 for typical-velocity impacts at the side of the helmet; this level would probably cause rotational head injury.

To realistically simulate oblique helmet impacts in crashes, the test headform should have a compliant scalp, and soft under-chin tissue, rather than being entirely made of metal. The impact velocity component V N normal to the test anvil should exceed 5 ms -1 to be typical of serious helmet impacts. The oblique impact tests currently in BS 6658 and EC Regulation 22 do not encourage the development of helmets that reduce peak rotational head accelerations, because the headforms lack the required features, V N is too low, and the pass/fail criteria are not expressed in terms of rotational acceleration.

The peak headform rotational acceleration was shown to be a function of three main parameters: the impact velocity component normal to the road, the friction coefficient between the shell and road, and the impact site/direction. . It was relatively insensitive to the tangential component of impact velocity. The main reason why potentially-injurious peak rotational accelerations of 20 krad s -2 occurred, for some impact sites when V N = 7.5 ms -1 , was the high linear head acceleration permitted by helmet test standards.


Marc Michel

Jérôme Bour

Julien Petersen

Claire Arnoult

Frank Ettingshausen

Christina Roth

David Ruch

Atmospheric Plasma Deposition: A new pathway in the design of Conducting Polymer-based Anodes for Hydrogen Fuel Cells

Keywords: Fuel Cell Electrode, Plasma Deposition Technology, MEA Preparation, Fuel Cell Technology, Assemblies

In this study, we explored thin films of nanofibrous functionalized conducting plasma Polyaniline (pPANI) with platinum deposited by an atmospheric plasma deposition process for the potential design of anodes for hydrogen fuel cell applications. We observed that the incorporation of such a polymer, characterized by both electronic and ionic conductivity, associated with a catalyst in a 3D porous network, could lead to an increased probability of the three-phase contact to occur. In this context, aniline was mixed with functionalized platinum nanoparticles and used as the precursor. The role of these functionalized nanoparticles was not only to act as the catalyst for fuel cell purposes, but also as nucleation sites promoting the formation of the nanofibrous pPANI thin film during the plasma polymerization. The morphology of the thin film was analyzed by scanning electron microscopy and the efficiency, in terms of energy conversion, was assessed in a single fuel cell test bench.

Introduction

Polymer electrolyte membrane fuel cells (PEMFC) are being considered as promising candidates for a number of commercial applications including their use in portable electronic devices and automobiles. One of the key challenges facing the widespread commercialization of PEM fuel cells is their high cost. The cost is largely associated with expensive materials such as a noble metal catalyst, and inefficiencies in their utilization due to a weak control on the so called triple phase boundary (TPB) [1], which is located in the membrane electrode assembly (MEA). To effectively utilize the catalyst, these regions, corresponding to the interfacial areas between the reactant, electrolyte and catalyst, must balance the ionic and electronic conducting media, allowing continuous pathways of the protons to the electrolyte and electrons to the electrode. These media influence the hydrophilic and hydrophobic properties of the composite material bringing a design challenge at the nanometer scale due the necessary integration of disparate materials.

In this context, it has been shown elsewhere that electrodes fabricated by incorporating a polymer, characterized by both ionic and electronic conductivity, in a so called Layer-by-Layer (LbL) [2] architecture can give a significant boost to the fuel cell performance, whereas less catalyst is employed. [3][4][5][6][7][8][9][10] Despite very promising results, this method of preparation can only be done in a batch process.

Inspired by these works, this paper aims at designing a new generation of multilayered anodes using Atmospheric Plasma Polymerization, which has become a promising technology for inline processing. Plasma polymerization is a versatile technique for the deposition of plasma polymer films with functional properties suitable for a wide range of applications such as protective coatings, biomedical materials, electronic and optical devices or adhesion promoters. [11][12][13] For example, plasma-induced polymerization has been extensively used for the preparation of conducting polymeric thin films including polyaniline (PANI). In film can be obtained and they explored its application in the fabrication of NO 2 gas sensor. [14] It is now well-established that plasma polymers are different from those fabricated by conventional wet synthetic methods: they are usually branched, highly cross-linked, insoluble, pinhole-free, and adhere well to most substrates.

Nevertheless, plasma-polymerization is mostly performed at low pressure, limiting its application to batch processes. In addition, low pressure plasma technology requires the use of costly vacuum pumping systems. [15] Among various atmospheric pressure plasmas, dielectric barrier discharges (DBDs) [16] have received a lot of attention due to the easy formation of a stable discharge and their scalability.

These discharges are usually generated in the space between two parallel electrodes, at least one of these being covered with a dielectric layer. [17][18] The functionality of the plasma deposited layer is mainly determined by the nature of the used vapour precursors or monomers. [19] These precursors or monomers, which lead to the coating formation, are diluted in a carrier gas, which is usually helium, argon, or nitrogen.

In this study, we explored nanofibrous pPANI thin films deposited using the atmospheric pressure dielectric barrier discharge process for the design of potential anodes for hydrogen fuel cell applications. We expect that the incorporation of such a polymer, characterized with electronic and ionic conductivity and associated with a catalyst in a 3D porous network could lead to an increased probability of the three-phase contact to occur. In this background aniline was mixed with functionalized platinum nanoparticles and used as the precursor. The role of these functionalized nanoparticles was not only to act as a catalyst for fuel cell purposes, but also as nucleation sites promoting the formation of the nanofibrous pPANI thin film during the plasma polymerization. The morphology of the thin film was analyzed by scanning To our knowledge, this paper is the first showing that electrodes for fuel cells can be manufactured by a sequential atmospheric pressure dielectric barrier discharge process of aniline and functionalized platinum.

Experimental section

Materials

Hexylamine, sodium borohydride (NaBH 4 ) 99%, 4-Aminothiophenol (mercaptoaniline), Aniline, and platinum (IV) chloride (PtCl 4 ) 99.99% were obtained from Aldrich and used as received. PtCl 4 was stored in a dessicator under vacuum.

Functionalization of Platinum nanoparticles

The synthesis and the functionalization of platinum nanoparticles were strongly inspired by a previous work. [20] Briefly, 150 mg PtCl 4 was dissolved in 35 mL of hexylamine in a 200 mL flask leading to the formation of an orange solution (solution 1). Then, 165 mg of 4-Aminothiophenol was dissolved in 15 mL of a methanol/hexylamine solution (50/50) (solution 2). Finally, 150 mg of NaBH 4 was dissolved in 20 mL of a water/methanol solution (50/50). 10 mL of hexylamine was added after complete dissolution of sodium borohydride (solution 3). Solution 3 was then poured into solution 1 under vigorous stirring at room temperature. The reaction mixture turned brown within a few seconds. After 30 s, solution 2 was added to the mixture. After 5 min, 100 mL of Milli-Q water was added. The resulting solution was then stirred for 20 min before being transferred into a separatory funnel. Just after phase separation, the water was removed and the organic phase was repeatedly washed with the same amount of water (100 mL). The volume of the dark brown organic phase was reduced to ca. 1-2 mL by rotary evaporation at 30 °C and transferred into a centrifuge tube

Plasma polymerization

In this study, the precursor used for the plasma polymerization consisted of 5 mL of functionalized platinum solution mixed with 5 mL of aniline (adjusted at pH=2).

The DBD discharge was generated between an earthed bottom aluminium plate and two highvoltage aluminium top plates protected by a 3.25-mm-thick glass plate. The gap between the electrodes was set to 2 mm. A commercial Nafion 117 membrane was used as the substrate and was positioned at the bottom electrode. The precursor was atomized by a TSI 3076 system at a constant atomization pressure of 2.10 5 Pa and injected into the gas before entering the plasma zone. The carrier gas composition, which was set by mass flow controllers, consisted of a N 2 /O 2 (97/3 vol. %) mixture with a gas flow of 20 L.min -1 . The gas mixture containing the precursor aerosol was injected into the plasma through a slit between the two top electrodes. The experiment was carried out at atmospheric pressure and at room temperature. During the experiment, the top electrode block moved over the sample at a constant speed (4 m.min -1 ) and in a range of 380 mm. The plasma coating consisted of 100 passes and lasted 10 min leading to the formation of a greenish film on the Nafion 117 membrane. For the coating process, plasma discharges were generated by an AC power supply with variable frequency. The frequency was set to 1.5 kHz and the voltage was a pseudo sinusoid as a function of time. Power was kept constant

Characterization Techniques

Materials characterization

The chemical structure of the coatings was evaluated by FTIR spectroscopy using a Bruker Optics Tensor 27 spectrometer. UV-Visible investigations were obtained by using a Perkin Elmer lambda 35 Spectrophotometer.

The pPANI-Pt fibers were investigated by X-ray powder diffraction using a diffractometer STOE STADI-P with Germanium monochromated Cu Kα radiation (λ= 1.54060 A) and a position-sensitive detector with 40° aperture. Average particle sizes are estimated by the Scherrer equation.

In addition, transmission electron microscopy (TEM) was employed to measure the Pt particle size and the particle distribution. A JEOL 3010 high resolution electron microscope equipped with LaB 6 filament and an acceleration voltage of 300 kV was used for the measurements. A small amount of the platinum dispersion was placed on a holey-carbon film covered Cu grid and left to evaporate.

The platinum loading on the pPANI was obtained by thermogravimetric analysis (TGA) with a Netzsch STA 490 under synthetic air (Linde(R), 80% N 2 , 20% O 2 ) atmosphere from 30 to 1000 °C with a heating rate of 5 °C min -1 .

The morphology of the LBL films was checked by scanning electron microscopy (SEM) using a JEOL JSM 6340 field-emission gun scanning electron microscope. The MEA was cut into tiny pieces and deposited onto a SEM holder before being observed.

Fuel cell testing

The MEA was tested in a single cell fuel cell test bench using a commercial cell hardware (Electrochem. Inc.) and was conditioned over night, until a steady state current was reached at a potential of 0.6 V. In hydrogen operation, 150 ml min -1 H 2 (N 5.0, Linde) are loaded with water in a humidifier (T=70 °C) and fed into the fuel cell anode. As cathode feed, high-purity oxygen was supplied (N 4.5, Linde) at 75 ml min -1 to the fuel cell pre-heated to 80 °C.

Current-voltage curves were obtained galvanostatically using an operated fuel cell test bench.

Results and discussion

Functionalization and characterization of Platinum nanoparticles

A major point for the design of the MEA by plasma polymerization was the control of the particles size after their functionalization with mercaptoaniline. It was expected in this work that the use of a mixture of aniline and functionalized platinum particles could help to promote the formation of aniline plasma polymer thanks to nucleation processes occurring during the coating. This functionalization was based on the reduction of platinum salt by a reducing agent (NaBH 4 ) in the presence of mercaptoaniline. [20] The synthesis described here used hexylamine as the intermediate capping agent. The final capping agent was introduced in the reaction mixture directly after reduction of the platinum salt. The rapid addition of mercaptoaniline after platinum reduction was consequently required in order to avoid irreversible particle flocculation phenomena.

Infrared spectroscopy was performed on the functionalized nanoparticles solution in order to check the presence of mercaptoaniline surrounding platinum nanoparticles (Figure 1). Absorption peaks above 3000 cm -1 are attributed to N-H stretching modes whereas the peak located at 1612 cm -1 is characteristic of NH 2 deformation. The C-N stretching mode is located at 1272 cm -1 . The aromatic ring exhibits vibrations at 1578 and 1480 cm -1 corresponding to (in-plane bending modes) and 815 cm -1 (C-H wagging vibration).

The morphology of the functionalized platinum was investigated by TEM. On the TEM pictures (Figure 2), one can see the presence of individual particles, suggesting that the functionalization had a positive effect on the dispersion: no aggregation was observed. The good dispersion can be explained by the fact that the presence of mercaptoaniline acted as a powerful capping agent. Note that this dispersion was kept stable for several months. The particles ranged in size from 1 to 4.5 nm with a maximum centred around 3 nm (Figure 3).

Manufacturing of a fuel cell anode using plasma polymerization

The underlying idea was to use the functionalized nanoparticles as nucleation agent able to promote the formation of pPANI. The plasma polymer film was directly deposited onto a commercial Nafion 117 membrane.

As mentioned previously, fuel cell electrodes require three essential properties: they should be electrically conductive, proton conductive and able to trigger a chemical reaction catalyzed by a metal. From the SEM images (Figure 4 A,B), one can see the presence of pPANI-Pt fibers with a diameter range of 50-70 nm. The film that exhibits a total thickness of about 2.5µm, presents a strong interdigitation of pPANI fibers, suggesting a dense 3D porous network. This morphology increases the probability of contacts between the strands, while substantial porosity also offers efficient mass transport. Electrical conductivities were measured on the coated film and were found to be 5. 10 -3 S. cm -1 .

Note that the experiment performed without Pt nanoparticles did not lead to the formation of fibers (Figure 4 C). This point suggests that the presence of functionalized Pt nanoparticles can act as nucleation sites, where pPANI can easily grow during the plasma process. The amount of Pt contained in the plasma polymer films can be obtained using thermogravimetric analysis (TGA) (data not shown) by completely oxidizing the fibers during analysis. TGA gives a Pt loading of 16 %, which represents 0.09 mg.cm -2 of Pt.

The X-ray diffraction (XRD) patterns collected for pPANI-Pt are shown in Figure 5. The diffraction peaks show that Pt nanoparticles are present in the face centered cubic (fcc) structure, as indicated by the characteristic peaks of ( 111), ( 200), ( 220) and (311). The average Pt particle size for pPANI-Pt catalysts was calculated from the broadening of the (220) diffraction peak using the Scherrer equation. The calculated mean Pt particles sizes for pPANI-Pt are 3.4 nm, which is consistent with the TEM observations. The FTIR spectrum of the pPANI-Pt thin film is presented in Figure 6. The characteristic peaks of emeraldine base form of pPANI were observed in the spectrum at 3260 cm -1 (N-H stretching with hydrogen); 3025 cm -1 (aromatic C-H stretching); 1532 cm -1 (C--C stretching of quinoid rings); 1480 cm -1 (C--C stretching vibration of benzenoid rings); and 1290 cm -1 (C-N stretching). The absorption band at 1125 cm -1 is assigned to N--Q--N bending vibration shift towards the lower wave number that corresponds to the pPANI-Pt emeraldine base. This result further supports the formation of pPANI-Pt using the plasma polymerization technique.

The pPANI-Pt thin film was characterized using a UV-visible spectrophotometer. In the spectrum (Figure 7), the characteristic bands of PANI (emeraldine base) were observed at 310, and 513 nm, due to the π-π* transition, polaron bands transition, and quinoid unit, respectively. These absorption bands exhibit both excitations of amine nitrogen of the benzenoid segments and imine nitrogen of the quinoid segments of pPANI-Pt. All these features clearly indicate the formation of pPANI by the plasma polymerization.

The performance of the prepared membranes onto the Nafion membrane was tested in PEM fuel cells (Figure 8). The film was used as anode and a standard carbon black (CB) (0.5 mg Pt respectively. The peak power density of the whole fuel cell was as high as 80 mW cm -2 , the obtained value remaining smaller than commercial fuel cells made of carbon black. This limited performance might be due to a low value of conductivity of the pPANI film.

Nevertheless, the Pt utilization, which is considered as one key affecting the practical and commercial prospects of fuel cell technology, can be obtained by dividing the peak power density by the Pt loading. In our case, the Pt utilization was found to be as high as 880 mW/mg Pt , which is a value comparable to those obtained by conventional electrode preparation methods of electrodes. This good value can be attributed to the nanoscale 3D organization of the plasma multilayered film adequate for balancing diverse requirements in the electrodes, and specifically to the significantly improved accessibility of Pt particles at the triple phase boundaries.

Conclusion

In this paper, we demonstrate the potential of atmospheric pressure plasma deposition for the design of conducting polymer-based anodes for hydrogen fuel cells. This work shows that the incorporation of functionalized platinum nanoparticles can act as nucleation sites able to promote the formation of polyaniline fibers during the plasma process. SEM images show that pPANI-Pt electrodes have a strong interdigitation and thus create a dense and extensive 3D network of fibers. This particular morphology increases the probability of contacts between the strands, while substantial porosity also offers promising mass transport. Hence, atmospheric pressure plasma represents an interesting new pathway for creating fast and efficient electrodes for hydrogen fuel cells.

  et al. have shown that nanofibrers of conducting polyaniline (PANI) thin

  the efficiency, in terms of energy conversion, was evaluated in a commercial single fuel cell test bench.

4 -

 4 Aminothiophenol dissolved in 7.5 mL of ethanol. After this solution was stirred overnight, a black solid residue was recovered by centrifugation. The supernatant presumably contained unreacted 4-Aminothiophenol and platinum complexes. The precipitate was washed with pure ethanol and centrifuged again. The fraction yielded to a dark brown solution, which remained stable for several weeks.

  and was set to 500 W, which corresponded to a power density (over the electrodes) close to 0.8 W.cm -2 .

  modes. Absorption corresponding to C-H bonds appears at 1172 cm-1

  based catalyst was used as cathode. H 2 and O 2 are used at the anode and the cathode,

Figure 1 .

 1 Figure caption.

Figure 2 .Figure 3 .Figure 4 .Figure 5 .Figure 6 .Figure 7 .Figure 8 .Figure 2 .Figure 4 .

 234567824 Figure 2. TEM pictures of functionalized Platinum nanoparticles with mercaptoaniline.

Wiley-VCH Fuel Cells


Hans-Juergen Meyer

Jochen Glaser

H.-Jürgen Meyer

Dr H.-J Meyer

An Improved Synthesis of Ternary Carbonates: The Examples

Keywords: Rare Earth, Carbonate Wiley-VCH ZAAC Rare earth, Carbonate, Metathesis, Crystal Structure, Luminescence

The ternary carbonate LiLa(CO 3 ) 2 , previously obtained under supercritical CO 2 pressure conditions was prepared by a new and efficient way of synthesis. A 2:1 molar mixture of Li 2 (CO 3 ) and LaCl 3 was reacted in a silica tube at 400 °C. The crystal structure of LiLa(CO 3 ) 2 was refined from X-ray single-crystal data and can be related with a caesium chloride type arrangement where Li + and La 3+ ions occupy the caesium positions and (CO 3 ) 2-ions occupy all cubic voids. Cerium-doped samples (LiLa(CO 3 ) 2 :Ce) exhibit blue luminescence. The mineral bastnäsite was synthesized by the same preparative approach, departing from a 1:1 molar mixture of LaF 3 and Li 2 (CO 3 ) being reacted at 450 °C.

Introduction

Carbonates of the rare earth (RE) elements with the expected formula RE 2 (CO 3 ) 3 remain a missing group of compounds, although the chemistry of corresponding carbodiimides has been recently developed, i.e. for the formula type RE 2 (CN 2 ) 3 with RE = Y, Ce -Lu [1]. The chemical relationship between carbodiimides and carbonates may be reflected by the synthesis of carbodiimide through ammonolysis reaction of a metal carbonate, and oppositely, by the decomposition of carbodiimide into carbonate in the presence of moist air or water, with a typical example being represented by the corresponding lithium compounds [2]:

Li 2 (CO 3 ) + 2 NH 3 = Li 2 (CN 2 ) + 3 H 2 O
Rare earth carbonates are known as hydrates, like La 2 (CO 3 ) 3 •(H 2 O) 8 [3] being crystallized from aqueous solutions of La(NO 3 ) 3 •6H 2 O and NaHCO 3 . The thermal decomposition of La 2 (CO 3 ) 3 •8H 2 O takes course over several steps. Unfortunately, the dioxide monocarbonate La 2 O 2 (CO 3 )•3H 2 O 4 is already formed below 150 °C to yield La 2 O 2 (CO 3 ) [5] near 400 °C [6].

Ammonolysis reactions of lanthanum carbonate hydrate have shown to yield lanthanum dioxide monocarbodiimide La 2 O 2 (CN 2 ) [7]. Therefore this preparative way has been considered inapplicable for the synthesis of anoxic carbodiimide compounds. [8] Ternary carbonate hydrate compounds having the general formula ARE(CO 3 ) 2 •xH 2 O (A = alkaline metal; RE = rare earth metal) have been already reported one-hundred years ago when their different solubilities were studied in attempt to develop a separation process for lanthanide elements. [9] A dehydration process of these compounds for the observation of non-aqueous compounds appears generally difficult, due to formation of dioxide monocarbonates

(RE 2 O 2 (CO 3 )).
A synthesis route for anhydrous ternary carbonates developed by Seidel and coworkers, departing from autoclave reactions of mixtures of alkali carbonates and rare earth oxalates, (RE = Y, La -Lu), but the crystal structures of these compounds are unknown. [10] Syntheses of ARE(CO 3 ) 2 compounds with A = Na, K and RE = Y, La -Lu were also reported following this way of synthesis. [11] The effort in operating equipment and performance related with the highpressure synthesis of anhydrous carbonate compounds is due the relatively low thermal decomposition temperatures, in case of LiLa(CO 3 ) 2 being obtained slightly above 400 °C. In this contribution a new and simple way of synthesis is described for the observation single crystalline LiLa(CO 3 ) 2 . In addition, the crystal structure and luminescence behaviour of a Cerium-doped sample of LiLa(CO 3 ) 2 are reported

Experimental Section

Syntheses

The synthesis of LiLa(CO 3 ) 2 was successfully performed by solid state metathesis reaction from LaCl 3 and Li 2 (CO 3 ) following a corresponding approach used for the synthesis of LiLa(CN 2 ) 2 [12] departing from LaCl 3 and Li 2 (CN 2 ). A 1:2 molar ratio of starting materials, LaCl 3 (ABCR, 99.9 %) and Li 2 (CO 3 ) (Alfa, ultrapure) was carefully ground in an agate mortar and loaded into a dry silica tube under inert gas conditions (glove box). The absence of contaminations with oxygen or water was crucial for the success of the reaction.

The argon filled silica tubes containing the reaction mixtures (total mass ca. 200 mg) were flame sealed under vacuum, transferred into a tube furnace and heated at 400 °C for 12 -24 hours.

After cooling to room temperature by the natural cooling rate of the furnace, crystalline mixtures of LiLa(CO 3 ) 2 and LiCl were obtained as reaction products by X-ray powder diffraction, without a detectable side-phase. Thus, the reaction may be described following the equation:

LaCl 3 + 2 Li 2 (CO 3 ) → LiLa(CO 3 ) 2 + 3 LiCl
Doped samples of LiLa(CO 3 ) 2 :Ce were synthesised by substituting 5 mol% of LaCl 3 by CeCl 3 and reacted under identical conditions as described before. Single-crystals of LiLa(CO 3 ) 2 were obtained when reacting the same LaCl 3 /Li 2 (CO 3 ) mixture with (≈ 2/3 equivalents) KCl in order to achieve an eutectic LiCl/KCl flux [13]. Products were washed with water to remove LiCl and the salt flux, respectively.

The same type of reactions was successfully performed for the synthesis of La(CO 3 )F (Bastnäsit). A corresponding 1:1 molar mixture of LaF 3 and Li 2 (CO 3 ) was reacted in a silica tube to yield La(CO 3 )F at 450 °C, according to the equation:

LaF 3 + Li 2 (CO 3 ) → La(CO 3 )F + 2 LiF

X-Ray Powder Diffraction

Diffraction data were collected on a Stoe StadiP X-ray powder diffractometer using Gemonochromatized Cu-K α1 radiation. The powder pattern of LiLa(CO 3 ) 2 was indexed monoclinically (from 27 reflections) yielding lattice parameters of a = 6.901

X-Ray Single-Crystal Diffraction

Transparent single-crystals of LiLa(CO 3 ) 2 were selected and fixed on the tip of glass fibers for a single-crystal diffraction measurement (Stoe IPDS diffractometer, graphite monochromatized Mo-K α radiation) at room temperature (Table 1). Intensities were corrected for Lorentz factors, polarisation and absorption effects. LiLa(CO 3 ) 2 was found to crystallize in the monoclinic space group I2/a. The structure refinement including anisotropic refinements of all atoms with the program SHELX [14] resulted in R1 = 0.0146, wR2 = 0.0366 for all measured X-ray intensities.

Crystallographic data are given in Table 2 and selected interatomic distances are provided in Table 3.

Photoluminescence Studies

A fluorescence spectrometer (Edinburgh Instrument, FS920, equipped with a 450 W Xe-lamp) was used for the measurement of excitation and emission spectra, all monitored by a cooled photomultiplier tube (Hamamatsu R2658, detection range: 300 -1010 nm) driven in a single photon counting mode.

Results and Discussion

The development of preparative routes for the synthesis of thermally moderately stable compounds is a great temptation in preparative chemistry because thermally labile solid state compounds may be regarded as the largest missing group compounds, with potential to exhibit new or enhanced functionalities. Several examples of solid state metathesis (SSM) reactions have been already successfully performed in a temperature region between 400 -600 °C. [15] A variety of rare earth carbodiimide compounds, i.e. those with the formula type RE 2 (CN 2 ) 3 were accomplished by SSM reactions departing from mixtures containing Li 2 (CN 2 ) and RECl 3 .

Exothermic reactions are obtained above 400 °C, sometimes near 500 °C, depending on the individual RECl 3 compound used, as can be monitored by thermo-analytic studies. The employment of a eutectic LiCl/KCl flux (melting point: ≈ 354 °C) has been advantageous in these reactions especially for single-crystal growth. Still lower reaction temperatures may be achieved by using an appropriate melt, or even a catalyst.

Binary rare earth carbonates that may be assigned with the expected formula RE 2 (CO 3 ) 2 remain unknown to date, whereas their carbodiimide counterparts have been developed, e.g. for the already mentioned formula type RE 2 (CN 2 ) 3 . Thermal stabilities of most rare earth carbodiimide compounds range up to 650 -750 °C and restrict any preparative attempt to operate below the decomposition temperature. The carbodiimide ion, which may be also assigned as dinitridocarbonate, is related with the (oxo-)carbonate ion by the same charge, by similar limited thermal stability, and by chemical transformation into each other mentioned in the introduction.

Thermal decomposition studies of LiLa(CO 3 ) 2 have already shown a decomposition to occur slightly above 400 °C. Therefore, all preparative attempts were restricted to this maximum temperature. The straight forward synthesis of a 2:3 molar mixture was successful for the preparation of pure LiLa(CO 3 ) 2 . No excess pressure was noticeable when opening the silica vessel containing the reaction product. A reaction performed with a 2:3 molar mixture of LaCl 3 and Li 2 (CO 3 ) targeting the composition La 2 (CO 3 ) 3 yielded LiLa(CO 3 ) 2 plus side phases.

Corresponding experiments departing from LaF 3 and Li 2 (CO 3 ) yielded La(CO 3 )F.

The counterpart of LiLa(CO 3 ) 2 in carbodiimide chemistry is LiLa(CN 2 ) 2 . The unknown crystal structure of LiLa(CO 3 ) 2 was refined on a single-crystal. The assignment of a monoclinic crystal system and the unit cell dimensions (Table 1) were consistent with the previously reported results. 10 The crystal structure of LiLa(CO 3 ) 2 , refined in this work for the first time can be considered to follow motif of the caesium chloride structure, in which Li + and La 3+ ions are arranged to occupy positions of Cs + ions and carbonate groups occupy all cubic voids of this arrangement (Figure 1). In fact, this idealized arrangement is clearly distorted as a result of the planarity and orientation of the carbonate group to allow for different nearest neighbour interactions with cations. Lithium ions are surrounded by four carbonate groups forming a distorted [LiO 4 ] tetrahedron with oxygen atoms of four carbonate groups (Figure 2). Lanthanum ions are surrounded by eight carbonate groups forming a distorted square antiprismatic arrangement with their carbon atoms (Figure 2). The arrangement of carbonate groups results in two bidentate nearest neighbor interactions and six monodentate interactions via oxygen atoms, resulting in the coordination number of ten at the lanthanum atom. The carbonate ion in the structure itself appears almost planar (Table 3).

The potential of LiLa(CO 3 ) 2 as a luminescent material was investigated by luminescence measurement on Ce 3+ doped samples of LiLa 1-x Ce x (CO 3 ) 2 with x = 0.05. The luminescence spectra of the Ce 3+ doped materials are governed by an interconfigurational transition between the ground state levels ( 2 F 5/2 and 2 F 7/2 ) of the [Xe]4f 1 configuration and the lowest crystal-field components of the [Xe]5d 1 configuration. This is a spin and parity allowed transition, resulting in high oscillator strength and a short decay constant. Due to the involvement of d-orbitals into chemical bonding, the energy of the crystal-field components of the [Xe]5d 1 configuration is a sensitive function of the chemical environment. Ce 3+ activated luminescent materials can thus show UV, blue, green, or red luminescence, whereby the energetic position of the absorption and emission bands are determined by the crystal-field strength and the covalent character of the lattice position, where Ce 3+ is localized. In case of LiLa(CO 3 ) 2 :Ce a blue emission is observed near 410 nm and the excitation spectrum monitored for the emission at 400 nm reveals two bands located at 315 and 285 nm (Figure 3).

A 2

 2 CO 3 and RE 2 (C 2 O 4 ) 3 •xH 2 O, under 2000 bar CO 2 pressure at 500 °C. Reactions with A = Li have yielded two monoclinic modifications of compounds having the general formula LiRE(CO 3 ) 2

  (2) Å, b = 7.842(2) Å, c = 8.432(2) Å, and β = 97.25(2)°. The powder pattern of La(CO 3 )F (Bastnäsit) was indexed hexagonally (from 10 reflections) to yield a = 7.173(4) Å, and c = 9.823(6) Å.

C( 1 Figure 1 .

 11 Figure Captions:

Figure 2 .Figure 3 .

 23 Figure 2. Arrangement of carbonate groups around lithium (left) and lanthanum (right) ions.

Table 1 .

 1 Crystal data and structure refinement for LiLa(CO 3 ) 2 .

	Page 7 of 11	ZAAC
	Empirical formula	LiLaC 2 O 6
	Formula weight	265.87
	Temperature	293(2) K
	Wavelength	0.71073 Å
	Crystal system	monoclinic
	Space group	I2/a
	Unit cell dimensions	a = 6.888(2) Å
		b = 7.853(1) Å
		c = 8.429(2) Å
		β = 97.24(3)°
	Volume	452.3(2) Å 3
	Z	4
	Density (calculated)	3.905 g/cm 3
	Absorption coefficient	9.393 mm -1
	F(000)	480
	Crystal size	0.2 0 x 0.06 x 0.06 mm -1
	Θ-range for data collection	3.95 to 25.91°.
	Index ranges	-8 ≤ h ≤ 8; -9 ≤ k ≤ 9; -10 ≤ l ≤ 10
	Reflections collected	2400
	Independent reflections	441 [R(int) = 0.0353]
	Completeness in 2Θ = 25.91°	99.3 %
	Refinement method	Full-matrix least-squares on F 2
	Data / restraints / parameters	441 / 0 / 47
	Goodness-of-fit on F 2	1.178
	Final R indices [I>2σ(I)]	R1 = 0.0144, wR2 = 0.0365
	R indices (all data)	R1 = 0.0146, wR2 = 0.0366
	Largest diff. peak and hole	0.705 and -0.770 e•Å 3
		Wiley-VCH

Table 2 .

 2 Atomic Coordinates and Equivalent Isotropic Displacement Parameters (Å 2 •10 3 )for LiLa(CO 3 ) 2 . U eq is defined as one third of the trace of the orthogonalized U ij tensor.

		x	y	z	U eq
	Li(1)	3/4	0.2515(11)	0	14(1)
	La(1)	1/4	0.1946(1)	0	8(1)
	O(1)	-0.0634(3)	0.3303(3)	-0.1241(3)	13(1)
	O(2)	0.4261(3)	-0.0851(3)	0.1207(2)	12(1)
	O(3)	0.1821(3)	0.4152(3)	0.2025(2)	14(1)
	C(1)	0.3963(4)	-0.0581(3)	0.2678(3)	10(1)

Table 3 .

 3 Selected Interatomic Distances (in Å) and O-C-O Angles (in °) for LiLa(CO 3 ) 2 .

	Li(1)-O(1)	1.862(4)	x2
	Li(1)-O(2)	1.975(6)	x2
	La(1)-O(1)	2.514(2)	x2
	La(1)-O(3)	2.517(2)	x2
	La(1)-O(3)	2.637(2)	x2
	La(1)-O(2)	2.649(2)	x2
	La(1)-O(2)	2.706(2)	x2

Acknowledgment

We are indebted to Dr. Helga Bettentrup and Prof. Dr. Thomas Jüstel from the University of Applied Sciences (FH Münster) in Steinfurt for recording luminescence data on LiLa(CO 3 ) 2 :Ce.