Journal of Occupational Accidents, 9 (1987) 199-211 Elsevier Science Publishers B.V., Amsterdam - Printed
in The Netherlands
Construction Site Workers Helmets A. GILCHRIST
and N.J. MILLS
Department of Metallurgy and Materials, University of Birmingham, Birmingham, B15 2TT (U.K.) (Received 4 February
1987; accepted 21 April 1987)
ABSTRACT Gilchrist, A. and Mills, N.J., 1987. Construction Accidents, 9: 199-211.
site workers helmets.
Journal of Occupational
Instrumented impact tests were carried out on various types of industrial safety helmets manufactured to BS5240, and force versus deflection graphs produced. A simple mathematical model was used to explain the results obtained when helmets are hit on the top. The model can then be used to predict the effect of modifying the shell stiffness or suspension stiffness. It also predicts the large force oscillations at 300 Hz caused by the striker exciting the resonant frequency of the shell mass/suspension spring system. It was concluded that while the helmets perform adequately during top impacts, the protection given at the sides, front and back of the helmet is so poor that a redesign is necessary.
BS5240 of the British Standards Institution (BSI) (1975) defines a safety helmet as “A helmet intended to protect primarily the upper part of a wearer’s head against a blow”. The standard states that the deceleration of a 5 kg striker with a 50 mm diameter hemispherical nose should not exceed 100 g when dropped from one meter onto the top of the helmet, and that the projected path of the striker should pass through the centre of mass of the headform. Safety helmets usually consist of two parts: the hard, smooth outer shell, constructed from ABS, polycarbonate or polyethylene; and the suspension cradle, made from a low density polyethylene moulding or a combination of a polyethylene headband and webbing straps over the top of the head. During an impact the shell will transmit some of the energy of the blow to the cradle, which will deform (viscoelastically or by yielding) and absorb some of the kinetic energy of the blow. In addition to impact considerations, the shell must be thick enough to resist penetration by a 3 kg conical striker falling 1 m onto sites within 50 mm of the top of the helmet, therefore requiring the helmet shell to be of a certain thick-
ness and the plastic to have a sufficiently high yield stress (penetration resistance is a function of yield stressx thickness). This has led to most manufacturers increasing the shell thickness at the top, which is against the spirit of clause 4.3 of BS5240 (BSI, 1975) which states that “the shell shall have as uniform a strength as possible and shall not be specially reinforced at any point”. However, it is permitted for the shell thickness to vary smoothly from the crown to the rim. The third mandatory requirement is that the shell should be flame resistant. The main constraints on design are (i) cost, as low cost will encourage companies to buy and use them, and (ii) comfort, as workers have to wear them for long periods of time. These twc constraints preclude the use of helmet,s designed to protect against major impacts ( such as motorcycle helmets) which are both too expensive, and far too uncomfortable for prolonged wearing. Since this research was performed BS5240 has been replaced (BSI, 1987). It is noticeable that the definition of a safety helmet now ends “. . . against a blow from above”, and the statement about “the shell shall have as uniform a strength. . .” has been removed. In other aspects the changes are minor and in particular the impact tests are unchanged. 2. METHODS
Three types of industrial safety helmets were chosen at random from the range available to the public (Table 1) . These were impacted using the method outlined in BS5240, and in addition were impacted on the sides, front and back, as research has shown (Procter and Rowland, 1986) that three times as many impacts are received on these sites as on the top. To do this the headform was tilted by 38” on a ball joint, TABLE
Helmets tested No.
1 2 3
Protector Safety U.K. Protector Safety U.K. Thetford Moulded Products
“ABS = acrylonitrile
Tuffmaster II Tuffmaster I Centurion 1100
ABS” ABS PE
PE = polyethylene.
324 291 290
1.9 2.1 1.6
3.9 3.7 5.3
which was then clamped before the striker fell vertically as before. These offcentre tests are not part of BS5240. The striker deceleration during impact was measured with an accelerometer mounted in the striker, and the force transmitted to the headform was measured with a quartz crystal load cell mounted under the headform. Both signals were captured using a Datalab 902 transient recorder, and subsequently analysed with a BBC model B microcomputer, to produce graphs of force vs. time; force vs. deflection, ‘and (Headform force - Striker force) vs. time. The latter demonstrates the ringing effect due to the impacter exciting the resonant frequency of the helmet. The programs used to analyse the results have been explained in an earlier publication (Gale and Mills, 1985). 2.2 Mathematical
Williams 1984 produced a simple one-dimensional model of Charpy impact tests. It shows how an impacting system can give rise to dynamical effects assuming that the object with mass m is impacted with an infinite mass A4 (which will not decelerate). The model uses two springs; K1, the elastic contact stiffness of the striker/specimen interface, and K,, the bending stiffness of the specimen (Fig. la). The model predicts that the force is given by
Fig. 1. Models used to simulate impact behaviour: (a) Williams, ing helmet behaviour; M, is the striker and M2 the shell mass.
(b) improved version for modell-
where P= force, V= velocity, t = time, and (I? = (K, + K2) /m where CL,is the natural frequency. It can be used to approximate the behaviour of a helmet during the first, couple of milliseconds of an impact. In this case, K1 is the contact stiffness of the bending of the shell, which was measured by supporting the rim of the helmet on a flat table on a load cell, and compression testing in an Instron testing machine with a flat anvil bearing on the crown of the helmet. The load was increased to 2 kN using a crosshead rate of 5 mm/min. K, is the slope of the graph of force vs. deflection which is linear to within 5%. The combined stiffness KT of the contact stiffnesss K, and the suspension system K2 in series was measured by supporting the helmet on a metal headform and compressing the crown of the helmet with a flat anvil. The relationship between the stiffnesses in this test is l/K,,. = l/K,
A flat anvil was used for convenience rather than the hemispherical one used during impact testing. Since the top part of these shells hardly bends during these tests, the shape of the anvil is immaterial. The thickest part of the shell (Table 1) is always at the crown, and the thinnest part is at the base. Since the bending stiffness is proportional to the cube of the shell thickness, the great majority of the deformation occurs in the lower half of the helmet. However other designs may deform significantly under the anvil. It was noticeable that the helmets sometimes buckled inwards under a hemispherical anvil when the contact point was 30’ from the top. Table 2 gives the suspension system type and the measured stiffnesses. Because the properties of plastics are time dependent these stiffnesses would be higher in an impact lasting 10 ms than in these compression tests lasting 0.5-2.5 min. The low density polyethylene (LDPE) cradle of helmet 3 has a more compliant design. The 6 arms meet at a lozenge shaped centre which can distort in shape. Consequently the helmet shell contacted the headform at a 1100 N load. Both bending stiffnesses are difficult to calculate theoretically, because of TABLE
and shell stiffness
Shell stiffness K, (kNm-I)
1 2 3
300 385 340
LDPE headband + 3 webbing straps LDPE headband and cradle LDPE headband and cradle
Stiffness K, (kNm-‘) 110 80 65
Fig. 2. The first part of the force vs. time graph calculated using the Williams parameters rn~0.3 kg, K,=300 kNm-‘,Kz= 110 kNm_‘, anda l-m drop height.
model for the
the thickness variation in the shell, and the complicated shape of the various parts that comprise the suspension system. The prediction of eqn. (1)) using typical values of K, and K2, is shown in Fig. 2. This graph does not closely model the results obtained from the helmet impact for two main reasons: (1) the helmet is not impacted with an infinite mass; (2) the helmet has a considerable damping effect associated with it, whereas the Williams model contains no mechanisms for energy loss. A better model consists of three masses, which can interact via springs and dampers when the masses are close enough together (Fig. lb). The motions are constrained to the vertical x-axis, so the motion of each mass has one degree of freedom. As the headform has an effectively infinite mass its motion is neglected. This leaves the motions of the striker and helmet shell to be determined. The striker is a rigid body, and the reference point on it is at the centre of its lower surface. The shell is a deformable body, and its reference point is its centre of mass. The reference point of the headform is its top point which is taken as the origin of the x-axis. The initial position of the centre of mass of the shell is at xp = L2, and the initial position of the top of the shell is at L, above the top of the headform. In the simulation the striker falls freely through a height h before impacting the shell surface. Table 3 gives the initial conditions. The forces between the bodies are calculated as follows: (a) the force fi between the striker and the shell is given by
of the simulation
ml, m, (kg) x1, x:, (m) v,, LI.,(ms ~‘) a,, a, (ms-‘)
ify, c 0 and
LL 0 0
where y1 = (x2 - L2) - (x, -L, ) and where K, is the shell stiffness spring constant and n1 the damper constant. The spring and damper only act when the striker is in contact with the top of the shell, when y1 > 0. (b) The force f2 between the shell and the headform is given by
if L2 > x2 > L:,
This means that the slope of the force deflection curve during ‘bedding-in’ is only one third of that after ‘bedding-in’. This approximates the curved forcedeflection curve found experimentally, a bedding-in distance L, - L,.,= 5 mm being used. Finally, allowance is made for ‘bottoming out’ when the striker hits the headform. The condition for this is x1 < 0, and a very high spring constant of KC,= 10 MNm-’ is used for the metal to metal contact. If bottoming out occurs, then the values of both fl and f2 from eqns. (1) and ( 2 ) are increased by -&x1. Newton’s second law is used to calculate the accelerations a: and a; of the striker and shell respectively, from (5) The * indicates that these are the ‘new’ values of the accelerations. The calculations are repeated at time intervals At, a typical value of At being 10 ps. Once the ‘new’ accelerations are known, the ‘new’ velocities and positions are found by the trapesium rule for numerical integration K=V,+~(a,+aT)At
where the unasterisked quantities are the ‘old’ values. The clock can then be advanced by an amount dt, so that the ‘new’ values become the ‘old’ values. The forces are then re-evaluated using eqns. (3) and (4) and the cycle of calculations repeated. Since the combination of spring and damper in parallel also occurs in the Voigt-Kelvin viscoelastic model, where it determines the retardation time s = n/K, it is of interest to evaluate the retardation times for the typical values used here. For the values Kl=300 kNm-‘, K,=llO kNm-‘, n,=lOO Nsm-‘, n2 =60 Nsm-‘, the retardation times are r1 =0.33 ms, r,=0.55 ms. From a knowledge of the response of the Voigt-Kelvin model to deformation at different rates (Mills, 1986)) we know that the springs dominate on a time scale much longer than T, and the dampers dominate on a time scale much shorter than r. It is not to be expected therefore that the model in Fig. lb will work for a slow compression test on a 1 minute time scale. Even in the impact lasting 10 ms the damping of the shell oscillations are relatively minor. This spring/damper model is only realistic at the particular time scale of the impact. For slow or ‘static’ loading the dampers have a negligible effect on the response. In reality the response of the low density polyethylene suspension being stretched, or a plastic shell buckling inwards ( Gale and Mills, 1985)) will show considerable hysteresis regardless of the time scale for loading and unloading. 3.RESULTS
3.1 Single impact testing Figure 3 shows a force vs. deflection curve for the top impact of a typical helmet (Tuffmaster I). BS5240 states that the deceleration of the striker should not exceed 100 g, i.e. the force on the striker should not exceed 5 kN when tested under the stated conditions. The only difference between these tests and the BS5240 test is that BS5240 requires the helmet to also be tested at - 10 a C and + 50’ C, whereas these are tested at room temperature (18’ C ) . The standard allows up to a minute between conditioning and testing. The authors have measured and modelled the temperature changes in ABS motorcycle helmets after conditioning (Gilchrist and Mills, 1987). When the same heat transfer parameters were used for a 2 mm thick ABS helmet, initially at - 10°C exposed to still air at 20°C on both sides, it was predicted that the shell was at 2 & 1 ‘C after 60 s; for a 4 mm thick shell after 60 s the temperature was - 3’ C at the surface and - 8 oC at the centre. Therefore the tests are not necessarily carried out at the nominal conditioning temperatures. As can be seen (Table 4)) all the helmets performed adequately, with a large margin of safety compared with the permitted maximum force of 5 kN. How-
Fig. 3. Typical striker force vs. deflection behaviour top with a 49 J blow. The rebound energy is 5 J.
of a construction
(2) struck on
TABLE 4 Standard
impacts of 49 J energy on the helmet crown
Tuffmaster II Tuffmaster I Centurion 100
1.90 2.80 2.27
ever the maximum deflections are high suggesting impact energy will cause ‘bottoming-out’. 3.2 Repeated
Figure 4 shows the force vs. deflection curve on the third impact on the top of the Tuffmaster II helmet. Although repeated impacts are not required by the standard, it is interesting to note that the maximum force increases progressively, and the third test caused the helmet to bottom out, as shown by the sharp peak, due to weakening of the shell and/or suspension system. It is probable that the suspension system has stretched and/or distorted so that the helmet sits lower on the headform, because the maximum deflection of 36 mm is similar to the values for new helmets. From this it can be concluded that helmets should be discarded after any serious impact, an advice which is given on the label attached to the helmet when new. 3.3 Off centre impacts Figure 5 shows a representative front impact with a 49 J blow. All results are of the same general type, transmitting a very low force up to a critical displace-
Fig. 4. Striker force vs. deflection
for the third top impact on helmet
Fig. 5. Headform force vs. striker deflection behaviour typical of an off centre impact of helmet 2). The transient recorder saturated at a force of 25 kN.
ment value, when the helmet shell hits the headform and transmits a very large force. This demonstrates that while the helmets provide a reasonable degree of protection from top impacts they are practically useless (for example, a 3 J side impact will cause over 5 kN to be transmitted to the headform) against side, front and back impacts. This is a good example of how manufacturers can design an article to fulfill the requirements of a standard, rather than designing
1 , ~-
0 0 L__ _
< .~. IO
_ _ _
3z r 82 y” e
_:--_:. _+*-.: ..-. 7 .._.& .---.
1 +- 7-----‘--- ..
__~ ..-. __ 30
_...i-..__ ..--5 -:~..--
Fig. 6. Predicted striker force versus deflection from the computer model with no damping. K, = 300. K,=llOkNm.-‘,h=lm. Fig. 7. As in Fig. 6 but with damper constants respectively.
of 100 and 60 Nsm
’ for the shell and suspension,
an article that will provide the best protection. By adding an energy absorbing foam around the sides of the helmet, or by changing the suspension system, a helmet could be produced to provide more complete impact protection, and research is proceeding here in this area. 3.4 Comparison
The predictions using the typical values of shell and suspension stiffness (300 and 110 kNm -l, respectively) are shown in Figs. 6 and 7. Figure 6 uses no damping, and therefore absorbs no energy, whereas Fig. 7 uses the empiricially chosen damper constants of 100 and 60 Nsm-’ for the shell and suspension system, respectively. Use of the dampers makes the prediction more realistic, especially in removing the force oscillations during unloading. The same spring and damper constants are then used in Fig. 8 to predict the striker force versus time; it is seen that this closely follows the experimental result for helmet (1) which has a webbing cradle. With other helmets (Fig. 3) having LDPE cradles the oscillations on impact can be more severe so lower values of the damper constants have to be used. 3.5 Force oscillations For all the helmets tested the measured force fi on the striker was not equal to the measured force f2 on the headform. Figure 9 shows a case where the peak value of fi exceeds the peak value of fi. If the force difference fl - f2 is plotted versus time it is found to be a damped sinusoidal oscillation, with a frequency
Fig. 8. Comparison of (a) an actual striker force vs. time graph (helmet 1) , with (b) the computer prediction using the same parameters as in Fig. 7.
23 t b Y 2
Fig. 9. Experimental variation of the striker force (solid curve) and headform force (dashed curve) with time for an impact where the peak headform force exceeds the peak striker force (helmet 2 ) .
of approximately 330 Hz. Such oscillations were observed with motorcycle helmets (Gale and Mills, 1985) where they are more severely damped by the polystyrene foam liner. The cause of the force oscillations is assumed to be the shell mass m2 vibrat-
ing on the two springs (suspension frequency of this system is
and shell) acting in parallel. The resonant
when the values K, = 385, K2 = 80 kNm- ‘, m2 = 0.3 kg are substituted the resonant frequency f for helmet 2 is predicted to be 200 Hz. The discrepancy between this and the measured frequency can partially be explained by the viscoelastic nature of the shell and suspension; the stiffnesses were measured on a 1 minute timescale and they will be considerably higher on the 3 ms time scale of the impact. 4. DISCUSSION
The ideal helmet would transmit a constant force to the head. There is some debate about the allowable maximum value F,,, for this force; BS5240 assumes it is 5 kN, but the motorcycle helmet standard BS6658 (BSI, 1985) assumes it is 15 kN. To date it has not been possible to obtain precise estimates of the forces that will cause concussion, permanent brain damage, or death. The two parameters for the design of a BS5240 helmet are then F,,, and the maximum deflection (or the distance between the crown of the helmet and the top of the headform). If the parameters are 5 kN and 40 mm respectively, then an ideal helmet (Fig. 10) could absorb a crown blow of energy 2005, whereas one with a linear loading curve would only absorb 100 J. In both cases there is no rebound of the striker. The actual response of BS5240 helmets is more complex than either of these idealisations, especially because of the induced oscillations of
5kN--_-___._-_____C_ I I 1
I I I I I I I
I I I
Fig. 10. Idealised force-displacement b - a linear response.
graphs: a -
with a constant
to the head;
the mass/spring system. Nevertheless the computer model can now be used to optimise the current type of design. It is important to realise that the force f1 acting on the striker is not equal to the force fi acting on the headform (Fig. 9). The point is that BS5240 requires that the deceleration of the striker remains below 100 g ( or the force fi remains below 5 kN) , whereas brain injuries will be determined by the magnitude (and possibly time dependence) of the force fi acting on the headform. It is possible that a peak in f2 could exceed 5 kN whilef, remained below 5 kN. Consequently the force transmitted to the headform should be measured. CONCLUSIONS
(1) Although the 3 helmets tested pass the crown impact criterion of BS5240, the shape of the force deflection curve is not ideal in optimising the energy absorbed. (2) It is possible to model the behaviour of a helmet using a simple spring and damper model producing results similar to those obtained during impact testing. (3) Oscillations of the mass and spring system mean that the force transmitted to the headform is not equal to that exerted on the striker. The force on the headform ought to be measured. (4) The three helmets provide little or virtually no protection for impacts on to the side, front or back of the head.
REFERENCES BSI, 1975. BS5240 General purpose industrial safety helmets. British Standards Institution, London, England. BSI, 1985. BS6658 Protective helmets for vehicle users. British Standards Institution, London, England. BSI, 1987. BS5240 Industrial safety helmets. Part I. Specification for construction and performance. British Standards Institution, London, England. Gale, A. and Mills, N.J., 1985. Effect of polystyrene foam liner density on motorcycle helmet shock absorption. Plastics and Rubber Proc. Appl., 5: 101-108. Gilchrist, A. and Mills, N.J., 1987. Fast fracture of rubber toughened thermoplastics used for the shells of motorcycle helmets. J. Mater. Sci., 22: 2397-2406. Mills, N.J., 1986. Plastics. E. Arnold, London, p. 132. Proctor, T.D. and Rowland, F.J., 1986. Development of standards for industrial safety helmets. J. Occup. Accid., 8: 181-191. Williams, J.G., 1984. Fracture Mechanics of Polymers. Ellis Horwood, Chichester, p. 238.