M Schnick
email: aschnick@mciron.mw.tu-dresden.de

G Wilhelm

M Lohse

U Füssel

A B Murphy

Three-dimensional modelling of arc behaviour and gas shield quality in tandem gas-metal arc welding using anti-phase pulse synchronization

Keywords: arc welding, tandem welding, GMAW, MHD arc modelling, arc interaction, gas shield quality

Three-dimensional modelling of arc behaviour and gas

shield quality in tandem gasmetal arc welding using anti-phase pulse synchronization M Schnick, G Wilhelm, M Lohse, U Füssel, a B

Introduction

Gas-metal arc welding (GMAW; also known at metal-inert-gas or metal-active-gas welding) is a process that is very widely used in industry to join metals. An electric arc is struck between a continuously-fed metal wire electrode (the anode, usually referred to as simply 'the wire') and the workpiece (the cathode).

Droplets of molten metal detach and fall to the workpiece, which is partially melted by the arc to form the weld pool. A shielding gas, usually argon or a mixture of argon and another gas, is fed around the wire in order to optimize the arc properties and to protect the hot areas of the weld pool from the ambient air. If air reaches this region, then weld imperfections can result. These include porosity of the weld, spatter (small droplets of molten metal), heat-tint oxidation (formation of coloured oxide layers on stainless steel that decrease corrosion resistance) and the deposition of fume (nano-and micro-particles formed from nucleation of metal vapour).

Tandem gas-metal arc welding (T-GMAW) has been developed in order to increase the productivity of the GMAW process. Compared to a single-wire process with similar welding parameters, it enables a doubling of the welding speed and sometimes an even greater increase in the deposition rate. T-GMAW is used in heavy steel industries, e.g. in shipbuilding or in pipeline production, for which high welding speeds are of strong economic importance.

In the T-GMAW process, two electrically-isolated wires are fed through a single welding torch, which controls the flow of shielding gas around the wires and into the arc. Thus two arcs are established into a single weld pool. Technical improvements of the welding torch design, the current waveforms, the power control and the arc synchronization in the last two decades have improved the process stability and applicability. However, compared to a single-wire process, there is an increased number of process parameters that have to be taken into account.

Optimization of the torch design and the welding parameters has so far been done experimentally. Recent publications demonstrate the strong influence of the angles of wires and the distance between the tips of the wires [START_REF] Hedegård | Enhanced prospects for tandem-MIG/MAG welding Document[END_REF][START_REF] Hedegård | The fundamental stability mechanism in tandem-MIG/MAG welding and how to perform implementation Document[END_REF][START_REF] Hedegård | High penetration tandem-MIG/MAG welding Document[END_REF][START_REF] Tanaka | Effects of torch configuration and welding current on weld bead formation in high speed tandem pulsed GMA welding of steel sheets Document[END_REF]. A large spacing between the wire tips promotes a large weld pool and a stable process.

A low spacing is used in order to obtain small weld pools. In this case, the arc attraction forces increase and synchronized pulsed arcs are indispensable in order to stabilize the system of arcs and to avoid waves in the weld pool that cause heavy activity in the built-in arc length control [START_REF] Goecke | Tandem MIG/MAG welding Svetsaren[END_REF]. The effects of weld pool waves can be minimized by short pulse lengths of about 1 ms [START_REF] Goecke | Tandem MIG/MAG welding Svetsaren[END_REF].

The most frequently-used approach is anti-phase pulse synchronization [START_REF] Trommer | Tandem wire process improves ship panel production[END_REF]. In this mode, the base current is present on one electrode while the second is in the pulse current phase and vice versa. However, abnormal increases of arc voltage and the arc interruptions occur sporadically, especially while one wire is in the base current phase and the other wire is at the beginning of the pulse current phase [START_REF] Tanaka | Occurrence of arc interference and interruption in tandem pulsed GMA welding Document[END_REF]. These effects occur particularly for welding with a shielding gas containing 10% or more CO 2 .

During T-GMAW with anti-phase pulse synchronization, strong interactions between the arcs can be observed, as shown in figure 1. The attractive forces deflect the arc that is in the base phase. The highspeed image demonstrates that the arc jet can be oriented perpendicular to the wire axis. During the pulse phase the arc is stabilized due to the strong plasma jet that driven by the magnetic pinch force (i.e. the j × B or Lorentz force) that results from the high current density at the top of the arc). This strong plasma jet can disturb the other arc that is in the base phase. Thus, the flow of the shielding gas will be affected and the flow field is expected to differ strongly from the shielding gas flow of single-wire GMAW processes, which has been investigated previously [START_REF] Dreher | Numerical optimization of gas metal arc welding torches using ANSYS CFX[END_REF]. Nevertheless, the gas supply of T-GMAW torches corresponds to that in standard GMAW torches, with the exception that an oval gas nozzle, rather than a circular nozzle, is used in T-GMAW in order to accommodate the two wires. The interactions between the arcs in T-GMAW have been investigated mainly experimentally [START_REF] Hedegård | Enhanced prospects for tandem-MIG/MAG welding Document[END_REF][START_REF] Hedegård | The fundamental stability mechanism in tandem-MIG/MAG welding and how to perform implementation Document[END_REF][START_REF] Hedegård | High penetration tandem-MIG/MAG welding Document[END_REF][START_REF] Tanaka | Effects of torch configuration and welding current on weld bead formation in high speed tandem pulsed GMA welding of steel sheets Document[END_REF][START_REF] Goecke | Tandem MIG/MAG welding Svetsaren[END_REF][START_REF] Trommer | Tandem wire process improves ship panel production[END_REF][START_REF] Tanaka | Occurrence of arc interference and interruption in tandem pulsed GMA welding Document[END_REF]. High speed cinematography was used to visualize the shape of the arc columns, but the images do not permit conclusions to be drawn about the arc temperature and flow field distributions [START_REF] Hedegård | High penetration tandem-MIG/MAG welding Document[END_REF][START_REF] Tanaka | Effects of torch configuration and welding current on weld bead formation in high speed tandem pulsed GMA welding of steel sheets Document[END_REF][START_REF] Goecke | Tandem MIG/MAG welding Svetsaren[END_REF][START_REF] Tanaka | Occurrence of arc interference and interruption in tandem pulsed GMA welding Document[END_REF][START_REF] Yudodibroto | Process stability analysis during tandem wire arc welding Document[END_REF]. It is possible to visualize the shielding gas flow field by Schlieren photography or particle image velocimetry (PIV) measurements, but while this has been done for standard single-wire GMAW [START_REF] Dreher | Numerical optimization of gas metal arc welding torches using ANSYS CFX[END_REF][START_REF] Schnick | Visualization and optimization of shielding gas flows by arc welding Document[END_REF], no results have been published for the T-GMAW process.

The first numerical investigations of tandem arc welding were done for two-electrode gas-tungsten arc welding [START_REF] Ogino | Numerical analysis on heat source characteristics of twoelectrodes TIG arc Document[END_REF]. These investigations were limited to a steady-state process with both arcs having the same current. The shielding gas flow was simplified by using a predefined inlet region, and diffusion of the surrounding atmosphere was neglected by assuming pure argon made up the whole fluid domain.

The work presented here is a major extension of previous investigations. A transient three-dimensional model of an anti-phase synchronized pulsed T-GMAW process is presented. While transient threedimensional arc models have been developed to investigate geometries of interest to welding [START_REF] Kaddani | Three-dimensional modelling of unsteady high-pressure arcs in argon[END_REF], plasma spraying [START_REF] Trelles | Multiscale finite element modeling of arc dynamics in a DC plasma torch Plasma[END_REF], plasma waste treatment [START_REF] Colombo | Three-dimensional time-dependent modeling of a DC transferred arc twin-torch system[END_REF] and circuit interruption [START_REF] Yang | Numerical analysis of the influence of splitter-plate erosion on an air arc in the quenching chamber of a low-voltage circuit breaker[END_REF], this is the first case in which such a model has been applied to a T-GMAW process. The model has been used to investigate the interactions between the arcs and the shielding gas flow in the process region. In particular, the causes of the deflection of the arc jet such as shown in figure 1, and the mixing of ambient air into the shielding gas, are investigated. Results are presented for pure argon shielding gas and for an argon-CO 2 mixture.

To allow the influence of the anti-phase synchronization to be more clearly determined, results are also given for a steady-state process with both arcs having the same current.

The paper starts with a description of the model, including assumptions, the governing equations, the simulation domain and the boundary conditions, in section 2. Results are presented and discussed in sections 3 and 4. Section 3 is concerned with the arc behaviour, and section 4 considers the effects of the arc behaviour on the gas shield quality. The results and conclusions are summarized in section 5.

Numerical model

The software package ANSYS CFX was used for a transient model of the tandem arc process. The standard equations of computational fluid dynamics (conservation of mass, momentum and energy), modified to take into account the electromagnetic and radiative phenomena that occur in thermal plasmas, were used [START_REF] Murphy | Thermal plasmas in gas mixtures[END_REF][START_REF] Gleizes | Thermal plasma modelling[END_REF][START_REF] Murphy | The effects of metal vapour in welding arcs[END_REF]. The changes required are the addition of resistive heating and radiative loss terms to the energy equation and a magnetic pinch term ( × j B ) to the momentum equation; j is the current density and B is the magnetic field. Two further equations are required: an equation for current continuity

( ) 0 σ ϕ ∇ ⋅ = ∇ ⋅ ∇ = j
, where is the electrical conductivity and is the electric potential, and an expression for the magnetic field = ∇ × B A , where the magnetic vector potential A is given by 2 0 µ ∇ = -A j , µ 0 being the permeability of free space. The effects of the sheaths, i.e. the cathode and anode fall zones, are simplified by using a mesh size of 0.1 mm in the plasma adjacent to boundaries of the wires the workpiece, as recommended in [START_REF] Lowke | LTE-diffusion approximation' for arc calculations[END_REF]. The ambient pressure is equal to atmospheric pressure.

The plasma is assumed to be in local thermodynamic equilibrium (LTE), as is usual for models of welding arcs. This has been shown to be a valid assumption in the main body of the arc, both experimentally [START_REF] Murphy | Laser-scattering measurement of temperature profiles of a free-burning arc[END_REF][START_REF] Murphy | Electron heating in the measurement of electron temperature by Thomson scattering: Are thermal plasmas thermal[END_REF] and theoretically [START_REF] Griem | Plasma Spectroscopy[END_REF]. Deviations from LTE can occur close to the cathode [START_REF] Haidar | Non-equilibrium modelling of transferred arcs[END_REF] and the anode [START_REF] Jenista | Numerical model of the anode region of high-current electric arcs[END_REF][START_REF] Yang | Anode attachment modes and their formation in a high intensity argon arc Plasma Sources[END_REF], and also in the fringes in the arc [START_REF] Cram | Departures from equilibrium in a free-burning argon arc[END_REF]. To take all the effects into account would require treating the transport and reactions of each plasma species separately. This would inevitably introduce inaccuracies because of the lack of available reaction data and uncertainties about appropriate calculation of thermodynamic and transport properties [START_REF] Rat | Treatment of non-equilibrium phenomena in thermal plasma flows[END_REF]28].

The thermodynamic and transport properties of the plasma were approximated using the pure gas properties of argon, CO 2 and air, taken from the work of Murphy and co-workers [START_REF] Murphy | Transport coefficients of argon, nitrogen, oxygen, argonnitrogen, and argon-oxygen plasmas Plasma[END_REF][START_REF] Murphy | Transport coefficients of air, argon-air, nitrogen-air, and oxygen-air plasmas Plasma[END_REF][START_REF] Tashiro | Prediction of energy source properties of free-burning arcs[END_REF] and then interpolating using the mixing laws of Gleizes et al [START_REF] Gleizes | Thermal plasma modelling[END_REF]. The mixture of argon to CO 2 is assumed to remain fixed at 82% argon and 18% CO 2 by volume, while the air fraction is variable. The presence of metal vapour and its influence on the arc properties and behaviour is neglected, as is the influence of the formation of metal droplets at the wires and their passage through the arc. Radiation is considered using the net emission coefficient model. For mixtures of argon and CO 2 , the net emission coefficient is calculated as a mole-fraction weighted average as recommended by Cressault et al [START_REF] Cressault | Thermal plasma properties in gas or gas-vapour mixtures[END_REF]. The net emission coefficients were calculated including self-absorption by assuming an isothermal cylinder of plasma of radius 1 mm, surrounded by a transparent environment. Data for argon were given by Cram [START_REF] Cram | Statistical evaluation of radiative power losses from thermal plasmas due to spectral lines[END_REF], and Cram's method was used to calculate data for carbon and oxygen.

The model implements the mixing of the shielding gas (argon or argon-CO 2 ) and air through turbulence as well as laminar diffusion. Turbulence was included using the shear stress transport turbulence model [START_REF] Menter | Two-equation eddy-viscosity turbulence models for engineering applications[END_REF]. This approach combines the K-model in regions near walls and the K-model in regions away from the walls. It has been rated the most accurate turbulence model compared to two other two-equation models and a one-equation model for a wide range of flows [START_REF] Bardina | Turbulence Modeling Validation, Testing, and Development[END_REF]. The diffusion of air into the shielding gas was modelled due to an additional transport equation for the air component. The laminar diffusion coefficient data was that given in [START_REF] Murphy | Transport coefficients of air, argon-air, nitrogen-air, and oxygen-air plasmas Plasma[END_REF] for diffusion of argon through air.

The computational domain includes a 180°mirror-symmetric geometry of the torch, the workpiece and the fluid (arc plasma) region, as show in figure 2. In order to take into account the influence of turbulence due to the radial boreholes in the gas distributor (see also [START_REF] Dreher | Numerical optimization of gas metal arc welding torches using ANSYS CFX[END_REF]), the shielding gas inlet was defined to be A non-uniform hexahedral mesh is used for all regions of the computational domain. The mesh is refined inside the torch and between the nozzle and the workpiece, in particular in the arc region, where the mesh size is 0.1 to 0.2 mm. The mesh contains approximately 1 900 000 nodes, of which about 1 400 000 nodes are in the fluid region.

The electric current of the arcs was defined by an anti-phase synchronized pulse regime, which is shown in figure 4. The pulse current is 442 A and the base current is 170 A. The time delay between the waveforms applied to the two wire electrodes is 2.4 ms. The pulse frequency is 252 Hz. The electric charge conservation equation is solved in all domains. Flux boundary conditions were used for defining the current profiles at the wire electrodes. The potential was set to 0 V at the base of the workpiece.

Conservation of the electric charge flux was enforced at all interfaces between the fluid and the solids. The workpiece cathode was modelled as a thermionic emitter, with an additional heat source term defined at the interface:

4 SB ( ) , c i i e w c q j U j T φ ϕ εσ = -- - & (1)
where U i is the ionization energy of argon, w is the work function of iron (4.5 V), is the emissivity (set to 0.5), T c is the cathode temperature and SB is the Stefan-Boltzmann constant. The electron current density j e is given by the Richardson equation: 2 exp ,

w e R c B c e j A T k T ϕ = - (2)
where A R is the Richardson constant, e is the electronic charge and k B is Boltzmann's constant, and j i is the ion current density, which is equated to the difference between the total current density and the electron current density.

At the wire anodes, the heat source due electron absorption at the surface was defined by 4 ,

a w SB a q j T ϕ εσ = - & (3)
where T a is the anode temperature. The heat source and sink terms due to vaporization, condensation and the electron velocity are neglected. The ohmic heating and the cooling due to the wire feed rate of 12.5 m/min are considered in the energy balance of the wires. The relative motion between the torch and the workpiece is neglected. Note that blackbody cooling terms in ( 1) and ( 3) are negligible compared to the other terms, so the use of a constant value of the emissivity will not cause a significant error.

The transient simulation was initialized by a steady-state solution for the tandem torch with 150 A in both wires, which allows the gas shield to be established. To investigate the transient arc and gas shield behaviour, a total simulation time of 16.1 s was used, with a constant time step of following discussion only refers to the last pulse of four that were simulated, extending from 12.0 to 16.1 s. The results for this pulse were nearly identical to those for the preceding pulse.

Arc behaviour and interactions

In the following discussion, 14 specified time points are considered in order to present a complete picture of the transient arc behaviour and its influence on the shielding gas flow. These are shown in figure 5.

They are non-uniformly arranged to allow detailed consideration of the effects of the increasing and decreasing currents. The left top picture of figure 6 shows the predicted arc temperature and flow map for a stationary tandem configuration in pure argon with a constant current of 150 A in both wires. It is obvious that the arcs interact and attract each other because of the superposition of the magnetic fields induced by the two arcs.

The columns of the two arcs meet above the workpiece, resulting in a high pressure area that causes upwards flow between the wire electrodes. The calculated maximum temperature is 18 000 K and the maximum plasma velocity is about 215 m/s. As for all the temperatures presented in this paper, we would expect the effects of metal vapour to decrease this value, since it has been shown that the increased radiative emission due to the presence of metal vapour emanating from the wire electrodes has this effect [START_REF] Murphy | The effects of metal vapour in welding arcs[END_REF][START_REF] Schnick | Metal vapour causes a central minimum in arc temperature in gas-metal arc welding through increased radiation emission[END_REF][START_REF] Schnick | Modelling of gasmetal arc welding taking into account metal vapour[END_REF].

The transient behaviour of the arcs is shown in the numbered images of figure 6; the numbers correspond to the times shown in figure 5. The image corresponding to time 1 shows the flow at the end of the base The waveform of the welding current (see figure 5) has two different downswing and upswing characteristics. In phase-shift mode 1, the upswing of the current in the left wire occurs while the right wire is in the pulse phase (figures 5 and 6: times 1 to 3). In contrast, in phase-shift mode 2, the upswing of the current in the right wire occurs while the current in the left wire is close to its base level (figures 5 and 6: times 10 to 12). This difference is found to have an important influence on the arcs.

In phase-shift mode 1, the maximum temperature of the arc with the increasing current is 1500 K lower and the plasma flow velocity is about 100 m/s less than in phase-shift mode 2. Further, in phase-shift mode 1, the arc properties are less steady in time and the radius of the arc near the workpiece is larger.

It was further found that the maximum temperature and plasma flow velocity of the left arc during its pulse phase (times 4 to 7) remain within narrow ranges of width 800 K and 100 m/s respectively. The maximum values were reached after the current of the right arc reached the base phase (time 5). In contrast, in phase-shift mode 2, the maximum values of temperature and plasma flow velocity were reached directly upon reaching the peak current (image 12), and then decreased slowly. Figure 7 shows the equivalent results to those given in figure 6, but for argon with 18% CO 2 . The top left figure again shows the predicted arc temperature and flow map for a stationary tandem configuration with a constant current of 150 A in both wires. In comparison to the results for pure argon, it is apparent that the plasma temperature is much higher and the plasma flow velocity is doubled. The model predicts a maximum arc temperature of 24 000 K and a maximum velocity of 450 m/s. The arc columns meet above the workpiece and there is an upwards flow between the wires, as was found for argon, but the arc are more constricted. The increase in plasma temperature and flow velocity when CO 2 is added to the shielding gas has been found in gas-tungsten arc welding arcs [START_REF] Tashiro | Prediction of energy source properties of free-burning arcs[END_REF] and also for gas-metal arc welding arcs in which the influence of metal vapour has been taken into account [START_REF] Schnick | Modelling of gasmetal arc welding taking into account metal vapour[END_REF]. The results obtained for the tandem configuration are consistent with those given in these references. As discussed in [START_REF] Tashiro | Prediction of energy source properties of free-burning arcs[END_REF], the increased arc constriction can be explained in terms of the high volumetric enthalpy of CO 2 . According to the 'thermal pinch' effect discussed by Tanaka and co-workers [START_REF] Tanaka | Predictions of weld pool profiles using plasma physics[END_REF][START_REF] Murphy | A computational investigation of the effectiveness of different shielding gas mixtures for arc welding[END_REF], the cross-sectional area through which heat flows is approximately inversely proportional to the volumetric enthalpy (the product of the stationary mass density and the specific enthalpy) averaged over the cross-sectional area, so the higher volumetric enthalpy decreases the cross-sectional area.. The higher current density in turn increases the resistive heating and the magnetic pinch force, thereby increasing the temperature and flow velocity.

The increased plasma temperature and plasma flow velocity are also apparent in the time-dependent behaviour in anti-phase synchronized pulsed T-GMAW (figure 7, times 1 to 14). Especially during the base phase, the arc is much more constricted than was found for pure argon (figure 6).

Of particular interest is the observation that the arc jet originating at the wire in the base phase is kinked (see, for example, the arc on the left wire at time 7 or the arc on the right wire at time 14). This behaviour is also clearly evident in the high-speed image shown in figure 1; the comparison with experiment will be discussed further below. Near the wire the jet is horizontal, but it becomes curved upwards as it moves towards the other wire. As the jet approaches the other wire, its direction is approximately parallel to the wire but in the opposite direction to that of the wire feed. A cold gas region occurs beneath the wire during the base phase. This cold gas forces the arc to kink upwards, which can lead to a stability problem, especially during the upswing of the wire current. It seems to be advantageous from this point of view for the upswing to occur while the other arc is in a low or base current phase. For example, the images corresponding to times 10 and 11, in which the right wire current is increasing from the base phase while the left wire is in the base phase, can be compared to those corresponding to times 1 and 2, in which the left wire current is increasing from the base phase while right wire is in the peak phase. The cold gas 'bubble' is much less prominent in the former case, and the flow velocity associated with the kinked arc is much lower.

The effects of arc instabilities have been investigated in depth by Mentel [START_REF] Mentel | Untersuchung der Wendelinstabilität an einem wandstabilisierten Wasserstoffbogen[END_REF]. He found that kink instabilities are promoted by convective cooling, and that heat conduction and resistive heating are stabilizing effects. This can be understood by considering images corresponding to times 6 and 7 of figure 7. The bubble of cold gas entering from the right cools the underside of the approximately horizontal arc, so that the conductive path is forced upwards. Heat conduction and resistive heating within the arc oppose this effect.

Figure 8 shows the volumetric enthalpy, electrical conductivity and thermal conductivity of argon and CO 2 at 1 atm. The volumetric enthalpy of CO 2 at the low temperatures (up to about 10 000 K) corresponding to those associated with the convective cooling is about 5 times larger than that of argon.

At the higher temperatures in the central region of the arc, where resistive heating and thermal conduction dominate (above 15 000 K), the thermal and electrical conductivities of argon and CO 2 are of similar magnitude. Thus, it can be seen that the addition of CO 2 increases the kinking due to the increased volume enthalpy, which leads to stronger convective cooling. The changes to electrical and the thermal conductivity are minor, so the stabilization effects are only weakly influenced.

The kink instability has a large influence on the wire temperature distribution. The instability leads to the flow of plasma towards and around the opposite wire, which preheats the wire before it is fed into the arc region. In addition to this direct heating, the electric conductivity of the wire is increased by the temperature higher temperature, therefore leading to increased resistive heating of the wire.

The power radiated from the arc plasma varies strongly with time. For the arcs in argon, it reaches a maximum of 2505 W when the two wires are at peak current, and a minimum of 866 W when the two wires are at base current. The total power input to the arcs has the same time dependence. At both times, the radiated power is 20% of the total energy input. Apart from the small fraction lost as heat in the wires, the balance of the power is transferred to the workpiece.

The addition of carbon dioxide to the shielding gas slightly increases the radiated power, to 2626 and 1084 W when the two wires are at peak and base current respectively. The increase is mainly attributable to the higher arc temperatures, rather greater net emission coefficient for the carbon dioxide. The net emission coefficient for the argon-CO 2 mixture is larger than that for argon for temperatures above about 19 000 K, and smaller for lower temperatures, but the difference is at most 10%. The total power input to the arcs increases by a larger factor, due the increased voltage associated with the greater arc constriction.

The radiated power is then 19% and 17% of the total power input at the two times.

It should be noted that the calculated arc voltages, which are in the range 14 to 19 V, do not include the sheath voltage, so the calculated input powers also neglect this factor. The measured arc voltage is 8 to 10 V higher; the discrepancy is partly due to the neglect of the sheath voltage, and partly due to the neglect of metal vapour in the calculation. It has been found that the presence of metal vapour leads to increased radiative losses in gas-metal arc welding, and this is sufficient to reduce the arc temperature and thereby increase the arc voltage [START_REF] Murphy | The effects of metal vapour in welding arcs[END_REF][START_REF] Schnick | Metal vapour causes a central minimum in arc temperature in gas-metal arc welding through increased radiation emission[END_REF][START_REF] Schnick | Modelling of gasmetal arc welding taking into account metal vapour[END_REF].

The Mach numbers are highest when both wires are near base current; although the velocities are lower at this stage because of the smaller Lorentz forces, the speed of sound is also smaller because of the lower temperature. The maximum Mach number is 0.234 for pure argon, and 0.248 for the argon-CO 2 case.

This indicates that the assumption of incompressible flow used in the model is a reasonable approximation.

The maximum Reynolds number is around 100, which means that the flow is laminar. Therefore the use of a turbulence model is in fact not necessary, and will not influence our results; this is in accordance with the findings for modelling of a single gas-metal arc [START_REF] Schnick | Metal vapour causes a central minimum in arc temperature in gas-metal arc welding through increased radiation emission[END_REF].

High-speed images of the tandem arcs at four different times are shown in figure 9. The images correspond to times 4, 6, 9 and 14, and can be compared to the temperature fields at the corresponding times shown in figure 7. The photographic images are not expected to exactly correspond to the temperature fields for a number of reasons. First, our calculations do not include the formation of droplets or metal vapour; both of these will affect the temperature, and droplet formation will alter the arc length.

Second, the images are integrals of the light emitted from the arcs in a given direction, while the temperature fields are shown for a cross-section through the arcs. Thirdly, radiation can also occur from the colder fringe regions of the arc due to radiative transfer processes [START_REF] Cram | Departures from equilibrium in a free-burning argon arc[END_REF][START_REF] Farmer | Local thermodynamic equilibrium in free-burning arcs in argon[END_REF]. Nevertheless, radiative emission from argon increases with temperature, so a reasonable correlation between the images and the calculated temperature fields is expected if the calculations are accurate. The image shown in figure 9(a) shows that both arcs are directed downwards when the two wires are near peak current. This in agreement with the temperature field at time 4 in figure 7. The bright region below the right arc is a consequence of a droplet being formed, leading to a much shorter arc than assumed in the calculations. Figure 9(b) shows the case when the left wire is at peak current (pulse phase) and the right wire at base current (base phase), while figure 9(d) is for the opposite polarity. The predicted upward kinking of the arc from the base-phase wire, shown at times 6 and 14, respectively, in figure 7, is apparent in both images. When both wires are near base current, the radiation is relatively weak owing to the low temperature. The upward kinking of the arc from the right wire predicted in calculations (time 9 of figure 7) is clearly visible in the image (figure 9(c)).

The predicted temperature fields show a strong correlation with the high-speed images, particularly when the intrinsic discrepancies are taken into account. This gives confidence that the most important processes, in particular those responsible for the interactions between the two arcs, have been included in the computational model.

Shielding gas flow and gas shield quality

The effect of the arc interactions on the gas shield quality is illustrated in figure 10. The air concentration is shown by the colour using a logarithmic scale, while the flow velocity is indicated by the colour of the arrows, using a linear scale. The length of the vectors is constant; shorter arrows result from vectors with a component perpendicular to the vertical plane. The arc temperature is indicated by the contour lines, which are the same as used in figure 7 (starting from 6000 K with a spacing of 2000 K).

The best gas shield quality (i.e. minimum air concentration in the vicinity of the wires and workpiece)

was found for the configuration with 150 A applied to both wires (top left). The nozzle design and the magnetic pinch forces resulting from the current flowing through the two arcs induce the flow of shielding gas between the arcs. The hot plasma is dispersed radially outwards and upwards. As a consequence, the flow field inside the torch and the gas nozzle is very different from the gas flow for single-wire torches. Hot gas is forced upwards between the wires and contact tubes while fresh cold gas flows downwards in the outer regions of the gas nozzle, particularly at the curved ends of the nozzle. The highest air concentration in the arc region was calculated to occur between the two wires. Air enters this region by diffusion from the sides parallel to the welding direction. However, the calculated air mass fraction is below 10 ppm, which is still very small and probably negligible in practice.

The use of anti-phase pulse synchronization decreases the gas shield quality considerably. The air mass fraction increases to more than 600 ppm in the process region. This level is sufficient to cause weld seam imperfections such as porosity, spatter and fume deposition.

The high level of contamination is caused by the convective flow in the shielding gas, and by diffusion and turbulent mixing. The air diffusion flux , D air m & is given by ,air eff air ,

x D m D C = - ∇ & (4)
where C air is the air mass fraction and D x eff is the effective diffusion coefficient of air in argon, given by eff air , Sc

x x t t D D µ ρ = + (5)
where is the density of air, D x air is the ordinary diffusion coefficient, which is a function of temperature and air concentration, µ t is the eddy viscosity and Sc t is the turbulent Schmidt number, which is set to 0.9.

It is found that the diffusion velocity is everywhere less than 0.2 m/s and everywhere at least an order of magnitude less than the flow velocity. This indicates that diffusion and turbulent mixing do not play a dominant role in determining the gas shield quality.

Figure 10 demonstrates further that the presence of air in the welding zone is mainly related to the flow associated with the base current arc. The horizontal flow associated with the upward-kinked arc leads to a horizontal intake of air into the arc. However, the flow field in the vertical plane through both wires that is given in figure 10 shows only relatively low flow velocities toward the base current arc, especially at the beginning of the base current phase. For a fuller picture of the flow field, it is necessary to consider other planes.

Figure 11 shows a plot of the flow vectors in three horizontal planes at the beginning of the base current phase of the left wire (time 6 in figure 5). The view is from above, and the contact tubes were rendered invisible in order to allow a full picture of the flow map. The absolute velocity value of the vectors is indicated by the colour scale, and the length of the vectors is constant. Shorter arrows result from vectors with a component perpendicular to the horizontal plane.

Consistent with the flow fields shown in figures 7 and 10, it is apparent that the flow field in the region above the arc is determined by the base current arc, while the peak current arc determines the flow close to the workpiece. The horizontal plane (a) is positioned 3 mm above the tips of the wires. The highest flow velocity is predicted to occur near the wire (1), which is in the pulse phase. This flow is induced by the horizontal base-current arc of the right wire. It is interesting to note that while this arc determines the flow field in the inner region, the flow in the outer region of the gas nozzle, especially the right side (2), continues to flow in the opposite direction, towards the right side. This flow is a result of the previous base phase of the left wire.

The horizontal plane (b) is positioned at the tips of the wires. The flow is mainly determined by the gas intake at the arc attachment to the wires. While the flow vectors near the left wire (in the pulse phase) are mainly directed downwards, the direction of the right arc (in the base phase) is horizontally to the left.

The main flow of the gas that is taken into this horizontal arc is coming from the sides (3). Thus, air from the outer nozzle region is drawn into the process region of the arc and the weld pool. The contaminating air remains in the gas shield flow for a long time, since it is transported upwards due to the arc deflection and then moves back and forth due to periodic changes corresponding to base and pulse current phases.

The horizontal plane (c) is positioned 2 mm above the workpiece and 3 mm below the tips of the wires.

The highest velocities are obtained near the pulse current arc (4). The flow field is pear-shaped. Below the right arc (5) the magnetic pinch force causes strong flow and kinking.

Conclusions

A three-dimensional MHD model of a tandem GMAW arc welding process has been developed for numerical studies of the interactions between the arcs and the resulting arc deflection, as well as to investigate air contamination within the gas shield. The model takes into account the complex geometry of the welding torch. An anti-phase pulse synchronization of the electric current was considered. The arc interactions and deflection were studied for pure argon and a mixture of argon with 18% CO 2 .

A complete model would take into account the production and transport of metal vapour, and its influence on the transport and radiative properties of the plasma. The influence of the formation of metal droplets would also be included. The emphasis here is, however, on the interactions between the two arcs, and high-speed photographs of the tandem arcs confirm that the essential features of these interactions are captured by the model.

The model demonstrates a strong deflection of the arc, especially during the base current phase. During the pulse current phase, the arc is stabilized due to the arc constriction at the wire. The arc deflection is increased by the addition of CO 2 in the shielding gas. This has been discussed in the context of kink instabilities, and has been attributed to the higher volumetric enthalpy of CO 2 .

Two different modes of phase shift were investigated: (1) the upswing of the current in the first arc while the second arc is still in the pulse phase and (2) the upswing of the first arc when the second arc has reached the base current phase. Phase shift mode 1 causes oscillations of the arc properties but also a large-radius arc above the workpiece. Phase shift mode 2 causes a smaller arc radius above the workpiece but steady arc temperature and flow velocity. The most distinct arc kinking was predicted for the phase shift mode 1.

The effects of arc interactions, especially the kinking of the base current arc, seem to play an important role in the introduction of air contaminations into the gas shield. Compared to GMAW tandem welding with two arcs of equal and constant current, the anti-phase pulse synchronization of the electric current causes a significant degrading of quality of the gas shield.

The deflection (kinking) of the arc in the base current phase is particularly responsible for high levels of air contamination in the process region between the arcs and the hot regions of the workpiece. The main flow of gas into the horizontal arc originates from the flat sides of the nozzle. This flow leads to ambient air from outer nozzle region being taken into the process region. This contaminating air remains in the gas shield flow for a long time since it is transported upwards due to the arc deflection and then oscillates due to periodic changes of base and pulse current phases of the wires.

Figure 1 .

 1 Figure 1. High-speed image of the interaction between the arcs during anti-phase synchronized pulsed T-GMAW.The arc on the left is in the pulse phase, and the arc on the right is in the base phase.

  above the base of the nozzle. A flow of 15 l/min argon (either pure or with 18% CO 2 ) at 310 K is used at each inlet. The fluid domain is bounded by the interfaces with the torch and with the workpiece, and by the symmetry plane and a surrounding open region consisting of pure air at 300 K and at ambient pressure. The fluid domain has dimensions 200 75 50 mm × × .

Figure 2 .

 2 Figure 2. Schematic showing the simulation domain of the two torches, the gas nozzle, the workpiece and the fluid region

Figure 3 .

 3 Figure 3. Geometrical parameters used for the tandem torch. Lengths are in millimetres.

Figure 4 .

 4 Figure 4. Welding current waveforms for the leading and the trailing wire.

Figure 5 .

 5 Figure 5. The 14 time points considered, relative to the currents in the two wire electrodes.

Figure 6 .

 6 Figure 6. Predicted arc and wire temperature and flow fields in the plane of symmetry for tandem welding in pure argon: stationary case of 150 A in both wires (left top) and transient behaviour at the times 1 to 14 indicated in figure 5. The leading wire is on the left.

  left (leading) wire and during the pulse phase of the right (trailing) wire. The arc temperature of the left arc is about 16 000 K. The flow direction of the left arc is nearly horizontal. The predicted maximum temperature of the right arc during its pulse phase is about 24 000 K. The flow of this arc is directed along the axis of the wire. During the upswing of the current in the left wire (times 1 to 4), the temperature and the plasma flow velocity increase rapidly and the flow direction shifts slowly back towards the direction of the wire axis. In general, the temperature and the plasma flow velocity change much more rapidly than does the direction of the plasma flow.

Figure 7 .

 7 Figure 7. Predicted arc and wire temperature and flow fields in the plane of symmetry for tandem welding in argon with 18% CO 2 : stationary case of 150 A in both wires (left top) and transient behaviour at the times 1 to 14 indicated in figure 5. The leading wire is on the left.

Figure 8 .

 8 Figure 8. Calculated thermal conductivity, electrical conductivity and volumetric enthalpy of argon and CO 2 at 1 atm, assuming LTE. The enthalpies are relative to the values at 300 K.

Figure 9 .

 9 Figure 9. High-speed images of the arcs for tandem welding in argon with 18% CO 2 , corresponding to the following time points in figure 5: (a) time 4 (both wires near peak current); (b) time 6 (left wire in pulse phase, right wire in base phase); time 9 (both wires near base current); (d) time 14 (left wire in base phase, right wire in peak phase).

Figure 10 .

 10 Figure 10. Predicted mole fraction of air and flow fields in the plane of symmetry for tandem welding in argon with 18% CO 2 : stationary case of 150 A in both wires (left top) and transient behaviour at the times 1 to 14 indicated in figure 5. The leading wire is on the left. Velocities are in m/s.

Figure 11 .

 11 Figure 11. Predicted flow field at time 6 in figure 5 for tandem welding with argon and 18% CO 2 at horizontal planes (a) 3 mm above the tips of the wires, (b) level with the tips of the wires and (c) 2 mm above the workpiece. The schematic diagram shows the point of view from which the horizontal planes are viewed.

On this basis, it appears that decreasing the arc deflection is a promising approach to improve the quality of the gas shield and to avoid weld seam imperfections such as porosity, spatter, heat-tint oxidation and fume deposits beside the weld seam, all of which are exacerbated by the presence of air in the process region. Several different approaches could be considered. The first, which has been clearly indicated by the work presented here, is to decrease the CO 2 concentration in the shielding gas. A second approach would be to decrease the arc length, in order to increase the stabilization of the arc due to constriction at the wire. Finally, the separation of the wires could be increased to decrease the interactions between the arcs.


J C F Millett
email: jeremy.millett@awe.co.uk

G Whiteman

S M Stirk

N K Bourne

Shear strength measurements in a shock loaded commercial silastomer

Keywords:

The shock induced shear strength of a commercial silastomer, trade name Sylgard 184 TM has been determined using laterally mounted manganin stress gauges. Shear strength has been observed to increase with increasing shock amplitude, in common with many other materials. Shear strength has also been observed to increase slightly behind the shock front as well. It is believed that a combination of polymer chain entanglement and cross linking between chains is responsible. Finally, a ramp on the leading edge of the lower amplitude stress traces has been observed. It has been suggested that this is due to shock induced collapse of free space between the polymer chains. Similar explanations have been used to explain the apparent non-linearity of the shock velocity with particle velocity at low shock amplitudes.

1.

Introduction

Polymers are widely used in industry, and thus their exposure to situations were high loading rates, including shock loading are becoming increasingly prevalent. The manufacture of explosives for military and industrial applications is a good example, where materials such as polychlorotrifluoroethylene (PCTFE) [START_REF] Anderson | Shock Compression of Condensed Matter -1991[END_REF] and hydroxyterminated polybutadiene (HTPB) [START_REF] Bohn | [END_REF]3] are used as binder phases. Epoxy resins have also been used as the matrix phase in fibre based composites [4,5], and as such their impact response is of interest to the aerospace industry where resistance to foreign object damage (FoD) and bird strike are a necessity. Interest in monolithic polymers such as neoprene (polychloroprene) [START_REF] Buland | 10th European conference on Earthquake Engineering[END_REF] has attracted attention as it has been used in the protection of buildings to earthquake damage.

In the past few years, a systematic effort has been made to understand how molecular structure affects the shock response of various polymers. One of the major aims of this programme was to investigate the variation of shock induced shear strength ( ) both with impact or longitudinal stress ( x ) and time behind the shock front itself. Such measurements are useful as they define the offset of the Hugoniot (longitudinal stress) from the hydrostatic pressure (P) via the relation, τ σ (

) 1
During shock loading, conditions of one-dimensional strain apply, and correspondingly, there must be a confining (or lateral) stress ( y,z ) orthogonal to the loading axis. Assuming that hydrostatic pressure is the average of all three components of stress, shear stress can be written in terms of these components of stress, thus,

y x σ σ τ - = 2 .
(

A number of techniques have been employed to determine this value, including inclined impact [START_REF] Zhou | [END_REF], subtracting the calculated hydrostatic pressure from the measured Hugoniot stress [8] and the self consistent technique, where a simple load and release experiment is compared to one where the initial conditions are the same, but instead of releasing, a second shock is launched into the already shocked target [9]. However, from equation 2, it can be seen that shear strength can also be measured using suitably orientated manganin stress gauges. This methodology has been used by ourselves and others on a variety of materials including metals and alloys [START_REF] Hopkins | Shock Compression of Condensed Matter[END_REF][START_REF] Millett | [END_REF][12], ceramics and silicate glasses [13][14][15] and polymers [16][17][18]. It can be seen that this is an invasive procedure, in that a manganin stress gauge must be introduced into a sectioned target assembly, and as such, it has been suggested that in doing so, the lateral stress recorded by the gauge is more a function of the layer of adhesive it sits in due to target assembly [19]. However, shock loading experiments below the Hugoniot Elastic Limit (HEL -the yield strength during one-dimensional strain) in ceramics and glasses using this technique show good agreement with the calculated elastic response, thus,

x σ ν ν τ - - = 1 2 1 2 , ( 3
)
where ν is the Poisson's ratio. Further, comparison of a number of techniques, including lateral gauges in tungsten heavy alloys (WHAs), also show a high level of agreement, hence lending confidence to the technique [20].

In the case of polymers, lateral stress gauges have been used to examine the role of side group size (from the main carbon-carbon) backbone has been investigated through the study of polyethylene (PE -hydrogen), polypropylene (PP -methyl) and polystyrene (PSbenzene) [18], showing that as the size of the side group increased, the level of strengthening increased. It was suggested that this increased the level of chain entanglement (tacticity) increased from PE to PS, thus increasing the necessary stress to achieve deformation. It was also suggested that both PP and PS showed evidence of increased crystallinity due to shock loading. Such a response has been observed in polyether etherketone (PEEK), where Millett et al. [21] showed a pronounced discolouration immediately behind the impact face during

Taylor impact tests, and by Rae et al. [22], who in repeating these experiments, were able to quantify this response. In all four materials, a significant increase in shear strength was also observed behind the shock front [17,18], again ascribed to a degree of tacticity between adjacent polymer chains. However, in the case of fluorinated polymers such as polytetrafluoroethylene (PTFE), PCTFE and polyvinylidene difluoride (PVDF), the situation is rather different [START_REF] Bourne | Shock Compression of Condensed Matter -2001[END_REF][START_REF] Bourne | [END_REF]. Although in common with most polymers studied, shear strength was shown to increase with longitudinal stress, shear strength itself was observed to be near constant with time behind the shock front. As fluorine replaced hydrogen, moving from PE to PTFE, strength was observed to shift from an increase behind the shock front, due to chain tangling to a more constant response. It was suggested that this was due to a two fold process; the large fluorine atoms reducing the tacticity of the chains, combined with the large electro-negativity of fluorine increasing the van de Waals forces between chains which will also affect strengthening mechanisms. Interestingly, it was noted that PVDF had the greatest strength of all three materials in this study. It was proposed that the alternating pairs of hydrogen and fluorine atoms along the chain would enhance tacticity, whilst the presence of those alternating pairs would also modify local charge density along the chain, thus allowing the creation of additional attractive forces.

We now turn our attention to a cross linked silicone rubber, marketed under the trade name Sylgard 184 TM , and referred hence forth as Sylgard. Although it is used as a potting compound in systems subjected to high pressure loading, and that it has been used as a binder phase in polymer bonded explosives [START_REF] Marsh | LASL Shock Hugoniot data[END_REF], it is perhaps surprising that very little information concerning its high strain-rate and shock response has reached the open literature. Winter et al. [START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF], in combination with previous data published in Marsh [START_REF] Marsh | LASL Shock Hugoniot data[END_REF] showed that the material displayed a linear relationship between shock velocity (U S ) and particle velocity (u p -the velocity of material flow behind the shock front due to momentum transfer). Silicone based rubbers have also been proposed by Be'ery and Rosenberg [START_REF] Be | Shock Compression of Condensed Matter[END_REF] as a potential encapsulation for manganin stress gauges to increase their pressure capability.

They suggested that this would be due to the reduction in shock induced conduction from pressure induced phase transformations in materials containing benzene rings such as epoxy based resins. Sander et al. [START_REF] Sander | Shock Compression of Condensed Matter -2005[END_REF], used an explosively generated shock of amplitude ca. 16

GPa and a maximum duration 30 ns to load Sylgard, monitoring its chemical response via mass spectrometer. Results showed that the polydimethylsiloxane polymer molecule decomposed into n-mers of dimethylsiloxane with n in the range 1 to 7. In addition, n-mers with missing methyl groups were also seen. The authors interpreted this as evidence of limited shock induced decomposition, suggesting that the lowest energy products had insufficient time to form. Dattelbaum et al. [START_REF] Dattelbaum | [END_REF], used diamond anvil techniques to compress Sylgard up to 10 GPa. They also used Raman spectroscopy, and were able to rule out pressure induced crystallization or phase transformations up to 6 GPa. Their results indicated close agreement between static and shock generated results. Hooper et al. [30], used molecular dynamics, based on simple quantum-chemistry based potentials to model a range of polymers, included Sylgard. Comparisons of predicted and experimentally determined compression, bulk modulus and shock response all showed good agreement.

In this paper, we use laterally mounted manganin stress gauges to measure the variation of shear strength in Sylgard, both with induced shock stress and time behind the shock front. We also re-examine the Hugoniot data generated by Winter and his colleagues [START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF] in light of past experience with other polymeric materials.

Experimental

All shots were performed using a 70 mm diameter, 3 m long single stage gas gun.

Impact stresses in the range 0.52 to 2.03 GPa were induced by impacting 6 mm flyer plates of either Dural (aluminium alloy 6082-T6) or copper, whose shock response is well documented [START_REF] Marsh | LASL Shock Hugoniot data[END_REF]. As Sylgard is a castable, two part resin, targets were manufactured as

follows. An accurately machined Dural ring, inner diameter ca. 65 mm by 10 mm deep was placed in the centre of a 2 mm thick driver plate of either Dural or copper, matched to the flyer. The driver plate was of diameter 120 mm, to allow the target assembly to be located accurately at the end of the gun. An accurately machined square bar of polymethylmethacrylate (PMMA) (10 x 10 x 80 mm 3 ) was used to support a manganin stress gauge (MicroMeasurements type J2M-SS-580SF-025) within the ring, such that the active element was ca. 4 mm from the surface of the driver plate. The active width of the gauge itself is ca. 240 µm, which gives a high degree of temporal resolution, allowing features on the rising part of the shock pulse to be identified. For example, using the elastic wave speed in Slygard (1.10 mm µs -1 discussed below in the materials section), the gauge would respond in ca. 220 ns. Gauge calibrations were according to the work of Rosenberg et al. [31], which accounts for the shape effects of lateral gauge results at low stresses. A schematic representation of the target assembly and gauge placement is shown below in figure 1. The elastic properties were determined using quartz ultrasonic transducers at a frequency of 5 MHz.

Materials

The Before moving onto the main subject of this paper, it is worth reviewing the existing Hugoniot data of Sylgard. We have taken data over the range of particle velocities of interest to this research from two sources; Marsh [START_REF] Marsh | LASL Shock Hugoniot data[END_REF] and Winter et al. [START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF]. In figure 3, we show the relationship between shock velocity and particle velocity.

Figure 3. Shock velocity versus particle velocity for Sylgard [START_REF] Marsh | LASL Shock Hugoniot data[END_REF][START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF].

Both sets of data appear to follow a standard line relationship of the form,

p S Su c U + = 0 , (4)
where c 0 and S are the shock parameters. Although there is an apparent difference between the two data sets, we believe that this is due to the degree of scatter present in both, and the fact that the particle velocity range is somewhat restricted. Therefore, we have followed the example of Winter et al. [START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF] and combined both sets of data to yielding values of c 0 and S of 1.22 mm µs -1 and 2.43 respectively. Observe that this value of c 0 (and indeed the separate values from figure 3) is higher than the ultrasonically measured value of c L . In metals, c 0 has been equated with the measured bulk sound speed (c B ), thus this discrepancy needs explanation. However, this behaviour has been observed in a number of other polymers, both by ourselves (see for example PEEK [17] and epoxy resins [16]) and in the compendium of polymer shock data published by Carter and Marsh [START_REF] Carter | Hugoniot equation of state of polymers[END_REF]. In semicrystalline polmers [START_REF] Bourne | [END_REF], it was suggested that this be due to a deviation below the accepted linear response at very low particle velocities by a transition from elastic to inelastic behaviour in the stronger crystalline phase. However, this cannot happen in amorphous or glassy polymers such as Sylgard, and at present, the issue remains unclear.

In figure 4, we show the combined Hugoniot data in stress -particle velocity space for Sylgard, taken from Marsh [START_REF] Marsh | LASL Shock Hugoniot data[END_REF] and Winter et al. [START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF].

with U S determined from equation 4 and the data presented in figure 3. The values of c 0 and S used are 1.22 mm µs -1 and 2.43 respectively, as discussed earlier. The degree of agreement between measured stress [START_REF] Winter | Shock Compression of Condensed Matter -2003[END_REF] and calculated pressure and measured pressure [START_REF] Marsh | LASL Shock Hugoniot data[END_REF] suggest, from equation 1 that the shock induced shear strength in Sylgard is low. We would point out that the hydrodynamic pressure discussed here is different from the hydrostatic pressure discussed earlier in equation 1. Shock loading is an adiabatic process, hence considerable heat is generated. This will result in the material trying to increase in volume due to thermal expansion, which will be resisted by the inertial confinement present during one-dimensional strain (all strain is accommodated along the loading axis, hence the material undergoes a volume reduction behind the shock front compared to ambient conditions). As a consequence, this resistance to thermal expansion will lead to an increase in pressure, hence resulting in the hydrodynamic pressure being greater than its hydrostatic counterpart.

4.

Results.

Representative lateral stress histories are presented in figure 5. Each trace is labeled with the longitudinal impact stress. Both traces have features in common; a slight overshoot after arrival of the main shock, and a noticeable decrease in lateral stress behind the shock arrival, which from equation 2 indicates a corresponding increase in shear strength. The overshoot is an electrical effect due to the fast rising nature of the shock pulse, and is discussed in greater detail elsewhere. Also observe that on the trace labeled 0.52 GPa, there is a low amplitude step feature on the signal rise, of duration ca. 180 ns. The first solution is that this is the result of the elastic response of the driver plate, in other words, the HEL of Dural. If true, this should correspond (approximately at least) with the wave separation of the elastic and plastic components of the shock in Dural over a distance of ca. 2 mm. From the elastic wave speed (6.40 mm µs -1 ) and the known shock response [START_REF] Hopkins | Shock Compression of Condensed Matter[END_REF] (U S =5.35+1.34u p , giving U S =5.54 mm µs -1 ), this results in a temporal separation of 50 ns, indicating that this solution is not correct. Therefore it would seem possible that this feature is a manifestation of the mechanical response of Sylgard.

The lateral stresses immediately behind the shock front have been used, along with the known longitudinal stresses to determine the shear strength of Sylgard. The results have been plotted against the corresponding longitudinal stress and the results are presented in figure 6.

Figure 6. Shear strength versus impact stress for Sylgard.

The straight line is the calculated elastic response according to,

x σ ν ν τ - - = 1 2 1 2 . ( 6
)
As was suggested in the materials section, the resultant shear strength during shock loading is very low (the shear strength in for example epoxy resin [19] over a similar stress range lies between 0.25 and 0.60 GPa).

5.

Discussion.

The main aim of this report is to investigate the shock induced shear strength in Sylgard.

From figure 5, it is clear that the material undergoes a significant reduction in lateral stress behind the shock front. Assuming that x is constant behind the shock front, then this implies, from equation 3, that the shear strength is increasing until releases enter the gauge location. Such a response has been seen in a number of materials, including face-centred cubic metals, where it was explained in terms of dislocation multiplication [START_REF] Millett | [END_REF], and more pertinently polymers such as PMMA [START_REF] Millett | [END_REF], PEEK [17], polycarbonate [34], epoxy [16] and polyethylene, polypropylene and polystyrene [18]. Clearly dislocation based mechanisms cannot be operating, especially as these materials represent a range of semi-crystalline (PE, PP, PEEK), amorphous glassy (PMMA, polycarbonate, PS) and cross-linked thermosets (epoxy). It should be noted that all of these polymers contain open structures which would allow a high degree of tangling (steric interference) between adjacent chains. Note from figure 2 that the PDMS chain consists of alternating silicon and oxygen atoms, with the silicon atoms linked to either methyl or phenyl groups. This gives an open structure which will again lead to a high degree of steric interference. In addition, the cross linking between chains will also provide an extra source of resistance to compression. In this respect, Slygard is behaving in a similar way to the cross linked epoxy, although the strengths of this material are significantly higher since it is below its glass transition temperature, whilst Sylgard is above.

In most lateral gauge experiments it is necessary to section the specimen in half to introduce the gauge to the correct location, before re-assembling with an adhesive layer.

Some reservations have been raised concerning this methodology [19], given that by its very nature, it is invasive. In a previous article, lateral stress gauge results in tungsten heavy alloy [20] have been shown to give the same results as those measured from pressure shear [35] or from differences between the Hugoniot and hydrostatic pressure [36]. Further, it has also been possible to relate changes in shear strength with both stress and pulse width to similar changes in shocked microstructure and spall strength [START_REF] Millett | [END_REF], hence there is a high degree of confidence in the lateral gauge response. However, with Sylgard, it has been possible to cast the uncured material around the gauge itself without the need for sectioning. The fact that Slygard behaves in a similar manner to many other polymers is a further indication of the validity of the lateral stress gauge technique to determine the shock induced shear strength.

The final feature requiring explanation is the small ramp noted on the lateral gauge trace labeled 0.52 GPa, shown in figure 5. We reproduce it below in figure 7, at a greater temporal resolution, below. We also include the rise times of two higher stress shots as well. Note that the ramp is also present at 0.98 GPa (although of lower duration and higher amplitude) and is absent by 1.21 GPa. The solution most likely lies in a non-linear Hugoniot at low particle velocities. In the materials section it was noted that c 0 , the zero particle velocity intercept of shock velocity lay significantly above the measured elastic wave speed, c L . It was suggested above that this may be due to a decrease in shock velocity below the linear response (see figure 3). In semi-crystalline polymers, it has been hypothesised that this behaviour may be due to elastic and inelastic deformation in the crystalline and amorphous phases respectively [START_REF] Bourne | [END_REF], but similar behaviour in amorphous polymers could not be explained in these terms.

Figure 7. The behaviour of the rising part of the lateral stress pulse in Sylgard.

However, examination of the PDMS molecule in figure 2 shows the possibility of significant free volume under ambient conditions. During the early stages of shock compression, this free volume has the potential to collapse under pressure, thereby ramping the stress pulse before the material can reach a denser state and shock up. Dattelbaum and her colleagues [START_REF] Dattelbaum | [END_REF] made similar observations concerning the non-linearity of U S -u p data at low particle velocities, which would also agree with the Hugoniot data discussed in section 3, where the zero particle velocity intercept, c 0 is greater than the measured longitudinal sound speed.

Conclusions

The response of the silicone based elastomer Sylgard 184™ to shock loading in terms of its lateral stress and shear strength. A re-examination of existing equation of state data, both from earlier work and from the literature gives an indication of a strong non-linearity at low particle velocities. A low amplitude ramp on the lower stress traces would appear to confirm this, and it has been hypothesised that the initially open structure of this cross-linked polymer would ramp the lower part of the shock. It has been noted that behind the shock front, lateral stress appears to decrease somewhat, suggesting a degree of hardening. This has been observed in a number of other polymers, and in common with those, it has been proposed that tangling between adjacent molecules (i.e. steric interference) is responsible.

Finally, it has been observed that the overall trend is for the shear strength to increase with increasing impact stress, in common with many other materials.

Figure 1 .

 1 Figure 1. Schematic diagram of target assembly and gauge placement. The right hand image

2 .

 2 figure 2.

Figure 2 .

 2 Figure 2. Simplified representation of the PDMS molecule and the crosslinking reaction.

Figure 4 .

 4 Figure 4. Hugoniot data for Sylgard.

Figure 5 .

 5 Figure 5. Representative lateral stress histories from Sylgard. Gauges are 4 mm from the

Figure 1 (Figure 2 (Figure 3 (Figure 4 (Figure 5 (Figure 6 (Figure 7 (

 1234567 Figure 1 (Figure1.tif)

Table 1 .

 1 1. Summary of Experimental Conditions.

	Dural	6	280	0.52
	Dural	6	486	0.98
	Copper	6	498	1.21
	Copper	6	719	2.03

Flyer Material Flyer Thickness (mm) Impact Velocity (m/s) Impact Stress (GPa)

Acknowledgments.

We would like to thank Pete Keightley, Mike Lowe, Nigel Park and Vince Durrant of AWE for their help and interest in this programme.

British Crown Copyright MoD/2010


Dr V A Khomchenko

L C J Pereira

J A Paixão

Substitution-driven structural and magnetic phase transitions in Bi 0.86 (La, Sm) 0.14 FeO 3 system

Keywords: 64.70.K-, 75.30.-m, 75.85.+t

Solid state synthesis and investigation of crystal structure and magnetic properties of Bi 0.86 La 0.14-x Sm x FeO 3 (0 x 0.14) ceramics were performed. It was found that a rhombohedral to orthorhombic phase transition took place in the series with decreasing average ionic radius of the substituting elements occupying the A-site of the ABO 3 perovskite. Magnetic properties of the compounds were shown to correlate with evolution of their structural state. Pure rhombohedral samples 0 x 0.06 were obtained in a mixed antiferromagnetic/weak ferromagnetic state. A small residual magnetization characteristic of the compounds was found to weakly depend on change of the chemical composition. Progressive increase of the residual magnetization was observed upon the rhombohedral to orthorhombic transition. Reasons of appearance of the weak ferromagnetism in orthorhombic and rhombohedral phases of the Bi 0.86 (La, Sm) 0.14 FeO 3 compounds were analyzed.

Introduction

Complex transition-metal oxides with the perovskite (ABO 3 ) structure are materials demonstrating a large variety of attractive physical phenomena such as hightemperature superconductivity or colossal magnetoresistance. Since 2000s, renewal of interest to magnetoelectric effect has given additional impulse to investigation of perovskite-type compounds. Among them, BiFeO 3 is the only material which is both magnetic and a strong ferroelectric (~100 µC/cm 2 ) at room temperature [1]. As a result it is the most extensively studied multiferroic system sometimes called "The Holy Grail" of multiferroicity [2]. In ferroelectric phase (i. e. below the Curie temperature of 1100 K), the compound possesses a rhombohedral R3c structure [3,4]. Symmetry of a high-temperature paraelectric phase is still a matter of intensive research [5][6][7]. Below the Neel point of 640 K, BiFeO 3 shows a long-range incommensurate cycloid-type antiferromagnetic structure with a period of ~620 Å along [110] h direction [8]. The spatial modulation prevents the observation of any net magnetization and the linear magnetoelectric effect. If the modulated magnetic structure in BiFeO 3 would be suppressed, the system would be weak ferromagnetic [9]. It was recognized, that the modulation can be eliminated by strain (either induced by substrate in epitaxial thin films or by doping) or by application of high magnetic field [2]. The spontaneous magnetization is known to arise in the substituted Bi 1-x Ln x FeO 3 (Ln= lanthanide) compounds [10][11][12], however, it is not quite clear how the appearance of the magnetization correlates with the substitution-induced structural changes (in other words, whether the antiferromagnetic-weak ferromagnetic transition coincides with a doping-driven structural/ferroelectric transformation).

It was recently proven that the crystal structure of Bi 1-x Ln x FeO 3 (Ln= Sm, Nd, La) compounds undergoes a ferroelectric-to-antiferroelectric structural phase transition at x~0.14-0.18 (the concentration boundary of the phase transformation shifts towards smaller x with decreasing ionic radius of substituting element) [13][14][15][16][17][18]. The antiferroelectric phase has the orthorhombic structure closely related to that characteristic of PbZrO 3 perovskite [19] and demonstrates a weak ferromagnetic behavior with a typical spontaneous magnetization of ~0.25 emu/g [18,20]. To reveal a relationship between the structural chemistry and magnetic properties of BiFeO 3 -based multiferroics it would be useful to follow evolution of the magnetization behavior near the morphotropic phase boundary. In case of Ln= Sm, the phase boundary corresponds to the x= 0.14 composition [13]. In Bi 1-x La x FeO 3 system, the same concentration is slightly smaller than the highest one for maintaining the initial noncentrosymmetric rhombohedral structure [21]. Taking into account these facts, we carried out the structural and magnetic characterization within co-substituted Bi 0.86 La 0.14-x Sm x FeO 3 (0 x 0.14) series. Such a substitution gives a compositional sweep allowing to estimate the evolution of the magnetic properties near the rhombohedral-to-orthorhombic phase boundary without varying a total concentration of rare-earth dopants in the host lattice.

Experimental

Polycrystalline samples Bi 0.86 La 0.14-x Sm x FeO 3 (0 x 0.14) were prepared by a conventional solid-state reaction method using the oxides Bi 2 O 3 , La 2 O 3 , Sm 2 O 3 , and Fe 2 O 3 taken in desired cation ratio. The synthesis was carried out in air in closed alumina crucibles at 960 °C for 8 h with heating/cooling rates of 5 °C/min. Phase analysis of the samples and investigation of their crystal structure were performed by X-ray diffraction (XRD) technique using automated Philips PW 1050/35 diffractometer with Co K radiation. XRD patterns were collected over an angular range 20° 2 100°w ith step of 0.02°and exposition intervals of 25 sec/step. The data were analyzed by the Rietveld method using the FullProf program [22]. Magnetic measurements were performed with a superconducting quantum interference device (SQUID) magnetometer (S700X, Cryogenic Ltd).

Results and Discussion

Rietveld refinement of XRD spectra obtained for the Bi 0.86 Sm 0.14 FeO 3 compound confirmed the expected coexistence [13][14][15] of the orthorhombic (~63 %) and rhombohedral (~37 %) phases (Fig. 1). Previous electron diffraction characterization of the orthorhombic phase showed that it should have the Pnam structure combining a complex (a -a -c + )/(a -a -c -) octahedral tilting (in Glazer's notation [23]) with antipolar displacements of A-site cations along [110/110] C directions of the parent cubic perovskite cell [16]. However, it was also proven that this NaNbO 3 -like tilting is not strong and/or sufficiently ordered to give any superlattice reflections in Xray or neutron diffraction experiments [16,20]. Accordingly, the space group Pbam (combines a -a -c 0 octahedral tilting with antipolar displacements of A-site cations along [110/110] C axis-the structure is characteristic of PbZrO 3 perovskite [19]) can be used in practice [16,20]. A minor rhombohedral phase (Fig. 1) retains the structural motif characteristic of parent BiFeO 3 [24]. With respect to ideal Pm 3 m perovskite, the ferroelectric R3c structure is obtained by antiphase a -a -a -octahedral tilting and offcenter ionic displacements along [111] C direction. Sketches of both Pbam and R3c structures are shown in insets of Fig. 1.

Lanthanum substitution in Bi 0.86 (La, Sm) 0.14 FeO 3 series results in a drastic decrease of the orthorhombic phase fraction: the x= 0.12 and x= 0.1 compounds contains ~64 % and ~82 % of the rhombohedral phase, respectively. For the x= 0.08 compound, the orthorhombic phase is almost completely suppressed: corresponding diffraction peaks are present only on a background level (accordingly, content of the phase should not exceed ~5 %). In case of 0 x 0.06 samples, no traces of the orthorhombic phase was reliably detected (Fig. 2 (a)). The results are consistent with the previous electron diffraction and high-resolution synchrotron X-ray diffraction experiments performed for Bi 1-x Sm x FeO 3 films and suggesting that the orthorhombic clusters can exist in the rhombohedral matrix in a broad compositional range down to x= 0.09-0.1 [14,15]. Average ionic radius of A-site ions in the threshold composition (and, hence, the Goldschmidt tolerance factor

) ( 2 O B O A r r r r t + + =
, where r A , r B and r O are the ionic radii of A, B and O ions, respectively, determining the degree of distortion in perovskites) is very close to that estimated for the Bi 0.86 La 0.06 Sm 0.08 FeO 3 compound.

This fact implies a similarity of nature of the substitution-induced structural evolution in Bi 1-x Ln x FeO 3 [13][14][15][16][17][18] and Bi 0.86 La 0.14-x Sm x FeO 3 systems and confirms that not only a concentration of polar-active Bi 3+ ions, but also a geometrical factor determines the doping-driven ferroelectric-to-antiferroelectric phase transformation in (Bi, Ln)FeO 3 .

Unit cell volumes of the rhombohedral and orthorhombic lattices in Bi 0.86 (La, Sm) 0.14 FeO 3 increase with decreasing Sm concentration without any pronounced anomalies on corresponding concentrational dependencies. For the two-phase compounds, the primitive cell volume of the orthorhombic lattice is about 0.6 % smaller than that obtained for the rhombohedral one. The step-like character of the primitive cell volume change (Fig. 2 (b)) suggests that the structural transformation takes place as a first-order transition. It is worth noting that increasing the temperature or duration of the heat treatment does not change the ratio between the rhombohedral and orthorhombic phases in Bi 0.86 La 0.14-x Sm x FeO 3 (0.08 x 0.14) samples, but leads to decomposition of the perovskite phases.

Before analyzing magnetic properties of the series under study, let us remind that parent BiFeO 3 demonstrates an antiferromagnet-weak ferromagnet phase transition at H C ~200 kOe [2]. The transition is associated with a magnetic-field induced removal of the initial spatially-modulated magnetic structure. The field-induced weak ferromagnetic phase possesses a magnetization of ~0.25 emu/g [2]. The induced magnetization coincides with the spontaneous one characteristic of weak ferromagnetic state of the orthorhombic antiferroelectric Bi 1-x Ln x FeO 3 compounds [18,20], thus suggesting a similar origin of the weak ferromagnetism in the rhombohedral and orthorhombic phases. A-site chemical substitution reduces the threshold magnetic field H C for cancelling the magnetic cycloid in the polar rhombohedral phase [11,21].

Taking into account these facts, the main experimental findings related to magnetic properties of the Bi 0.86 La 0.14-x Sm x FeO 3 compounds (Fig. 3) can be summarized as follows:

1. Rhombohedral samples, structurally-homogeneous in approximation of the integrated picture given by powder XRD, demonstrate nonzero residual magnetization (M r ~0.03-0.035 emu/g) incompatible with a pure antiferromagnetic state.

Comparing the residual magnetization with the value characteristic of spontaneous or

field-induced weak ferromagnetic state in BiFeO 3 -based compounds suggests that the rhombohedral samples contain 12-14 % of a weak ferromagnetic phase. Thus, compositional dependence of the residual magnetization correlates with a structural phase evolution of the Bi 0.86 La 0.14-x Sm x FeO 3 compounds. Rhombohedral samples possess the main antiferromagnetic state. Significant increase of the net magnetization driven by change of the average ionic radius of A-site substituting elements follows the increasing of orthorhombic phase content (see inset in Fig. 3).

Let us consider the origin of the spontaneous magnetization in orthorhombic phase of the Bi 0.86 (La, Sm) 0.14 FeO 3 compounds. First of all, it is necessary to emphasize that increasing the net magnetization taking place in the system with increasing samarium concentration is not related to magnetic contribution of the magneticallyactive Sm 3+ ions. Indeed, rare-earth (RE) subsystem in REFeO 3 ferrites orders magnetically only below T~5 K [25], so magnetic properties of Bi 0.86 La 0.14-x Sm x FeO 3 samples at room temperature must be completely determined by iron sublattice. In the ion is surrounded by six nearest neighbors with antiparallel magnetic moments.

Accordingly, the only reason of appearance of the spontaneous magnetization in orthorhombic phase of Bi 0.86 (La, Sm) 0.14 FeO 3 samples is a canting of antiferromagnetic sublattices. In fact, there are two mechanisms giving rise to a canting of the antiferromagnetic alignment of spins: the single ion magnetic anisotropy and the antisymmetric spin-spin interaction of the form D•(S 1 ×S 2 ), where D is the so-called Dzyaloshinsky vector (the microscopic origin of this term is attributed to the anisotropic superexchange interaction including spin-orbital coupling) [29,30]. For compounds with a high Neel temperature the latter mechanism is known to be more important (the Dzyaloshinsky-Moriya interaction is proportional to the exchange interaction, while the anisotropy energy of single spin nature, which comes from the spin-orbit coupling under the crystalline electric field, does not depend on the superexchange mechanism) [29,30].

In case of rhombohedral samples, existing net magnetization must be attributed to the suppression of the initial cycloid-type modulated magnetic structure. Indeed, symmetry of the space group R3c permits the existence of the Dzyaloshinsky-Moriya interaction [9], however the spatial magnetic modulation results in a periodic variation of the canting angle between the magnetic sublattices, so the magnetization averaged over the cycloid period becomes zero. Accordingly, destruction of the spin cycloid is a necessary condition for the appearance of the spontaneous magnetization in the parent structural phase. Taking into account that decreasing the average ionic radius of Asite ions in the rhombohedral Bi 0.86 (La, Sm) 0.14 FeO 3 samples effectively influences the major antiferromagnetic phase to reduce the field of the induced magnetic transition, but has no appreciable effect on residual magnetization, one could suggest that geometrical factor cannot be the main one directly responsible for appearance of the spontaneous magnetization in the rhombohedral phase. Apparently, quantity of the polar-inactive dopants favoring the suppression of dipole ordering is the more important factor controlling stability of the magnetic structure modulation (cycloidal modulation in bismuth ferrite is a result of magnetoelectric interaction [2]) and determining magnetic properties of the rhombohedral compounds. The assumption is consistent with behavior of the concentrational dependencies of the magnetization reported for the low-doped Bi 1-x Ln x FeO 3 compounds, when the residual magnetization gradually increases with increasing the lanthanide content [31]. Local deviation from a homogeneous distribution of the substituting elements typical of perovskite solid solutions [32] or/and lattice defects might be the driving force of the local suppression of the initial spatial modulation in the rhombohedral phase to induce appearance of the small magnetization [18].

Conclusions

In conclusion, investigation of crystal structure and magnetic properties of Bi 0.86 La 0.14-x Sm x FeO 3 (0 x 0.14) ceramics was carried out with XRD and SQUID- magnetometry techniques. It was shown that 0 x 0.06 samples retain the rhombohedral structure typical of parent BiFeO 3 . Further increasing Sm content leads to appearance of the orthorhombic phase possessing the structure closely related to that characteristic of

3 . 4 . 5 .

 345 The rhombohedral samples demonstrate a metamagnetic behavior similar to that observed in Bi 1-x Ln x FeO 3 compounds upon the field-induced suppression of the magnetic cycloid[11,18,21]. Threshold field of the phase transition decreases with increasing Sm concentration: a clear deviation from linearity expected for antiferromagnets is observed above H C 35 kOe for x= 0.02, H C 30 kOe for x= 0.04, and H C 25 kOe for x= 0.06 (with increasing magnetic field). Within concentrational range of the pure rhombohedral compounds (0 x 0.06), the percentage of the weak ferromagnetic phase weakly depends on the chemical composition (the samples have almost the same residual magnetization). Residual magnetization rapidly increases with increasing orthorhombic phase content reaching a value of 0.2 emu/g for x= 0.14 sample containing ~63 % of orthorhombic phase.

Bi 1 -e

 1 x Ln x FeO 3 compounds, iron has the valence state Fe 3+ and the high-[27,28], so ordered spins should create a G-type antiferromagnetic structure, in which each Fe3+

PbZrO 3

 3 antiferroelectric. Both structural phases coexist in a broad compositional range up to x= 0.14. Pure rhombohedral samples were found to possess a small residual magnetization, whose value very weakly depends on change of the average ionic radius of the substituting elements occupying Asite of the perovskite. Comparison of the magnetization with that expected for weak ferromagnetic state in BiFeO 3 -based

Figures 1 . 3 .

 13 Figures

Figure 3

 3 Figure 3

compounds suggests that the major magnetic phase (~86-88 %) of the rhombohedral samples remains antiferromagnetic. The compounds possessing the main rhombohedral phase demonstrate a metamagnetic behavior; the threshold field of the magnetic transformation decreases with increasing Sm concentration. Progressive increase of the net magnetization correlating with the structural evolution takes place with increasing orthorhombic phase content. Reasons explaining existence of the spontaneous magnetization in orthorhombic and rhombohedral phases of the Bi 0.86 (La, Sm) 0.14 FeO 3 compounds were discussed. V. A. K. is grateful to the Foundation for Science and Technology of Portugal (FCT) for financial support through the "Ciência 2008" program.


Phonethip Banouvong

Pierre Levi

Mothers and Children Nutritional Status and Food Habits in the City of Vientiane as Related to the Level of Urbanization Eric Bénéfice IRD Laos in collaboration with

Background Transitions

• Demographic transition: shift from high fertility and mortality to low fertility and mortality (Warren, 1929;Landry, 1982) • Epidemiological transition: shift from a pattern of high prevalence of infectious and malnutrition, to one of high prevalence of chronic and degenerative disease associated with urban-industrial lifestyles [START_REF] Omran | The epidemiologic transition: a theory of the epidemiology of population change[END_REF] • Nutritional transition: shifts on diet (refined food, low fiber content) and physical activity (lower levels) patterns [START_REF] Popkin | The nutrition transition: worldwide obesity dynamics and their determinants[END_REF]

Asia

• Asian countries are undergoing a rapid pace of urbanization, industrialization, technological changes that would impact their lifestyle -Global improvement of health status (better access to health facilities, better knowledge…) -Increase in non communicable diseases (CVD, diabetes, metabolic abnormalities, obesity…)

• Many studies from China, Pakistan, India, Philippines…) • Thailand [START_REF] Vallée | Urbanisation et santé à Vientiane (Laos): Les disparités spatiales de santé dans la ville[END_REF] Mechanisms [START_REF] Popkin | The nutrition transition: worldwide obesity dynamics and their determinants[END_REF] Dietary habits and urbanization

• Significant differences in contribution of food groups to energy content (General linear model of analysis of variance):

-Cereal: p<0.06 (area 3>2>1) -Meat: p=0.02 (area 1> 2 and 3) -Fish: p=0.03 (area 3>2>1)

• Contribution to protein content Greater intake diversity in strata 1 and 2 than in strata 3 (p< 0.003)

Nutritional status of under 3 yo children

• Overall:

-Wasting 7.5% (95% CI: 4.7~11.3%) -Stunting 13.9% (95% CI: 10.0~18.6%) -Underweight 15.0% (95% CI: 11.0~19.8%)

• No effect of urbanization level on either wasting or stunting Linear growth determinants

Commentaries

• A transition in food intakes pattern has been observed elsewhere in PDR Lao:

among "traditional ethnic groups" (Jutta [START_REF] Krahn | The dynamics of dietary change of transitional food systems in tropical forest areas of Southeast Asia[END_REF] -At the national level (Kennedy et al, 2011) • However, in Vientiane data for comparison purposes are lacking

-Comparability and accuracy of food surveys is an issue

Highlights: 4 points 1. The pace of nutritional transition is slower in PDR Lao compared with other Asian countries 2. Urban households still retained many positive traits of the traditional food habits: importance of fish, vegetables, fruits consumption… 3. Until now, the transition process is associated with an improvement in nutritional status 4. However, in certain segments of the population, situation could deteriorate quickly

economic growth, technological changes for work, leisure, & food processing

	Stage 3 Receding Famine	Stage 4 Degenerative Disease	Stage 5 Behavioral Change
	• starchy, low variety,	• increased fat, sugar,	• reduced fat, increased
	low fat, high fiber	processed foods	fruit, veg,CHO,fiber
	• labor-intensive	• shift in technology of	• replace sedentarianism
	work/leisure	work and leisure	with purposeful changes
			in recreation, other activity
	Nutrient deficiencies, weaning disease, stunting	obesity emerges, bone density problems	reduced body fatness, improved bone health
	Slow mortality decline	Accelerated life expectancy, shift to increased NR-NCD,	Extended health aging,
		increased disability period	reduced NR-NCD

5

Urbanization,

Transition in PDR Lao?

	PDR Lao is not as urbanized as other SE Asian
	countries (ex Bangkok ›10 million inhabitants
	vs Vientiane 300 000). But:

• Important economic growth (8 to 10% yearly) • Infrastructures (roads, dams, bridges) • Urban development (not only Vientiane) • Improvement of medical facilities • The part of agriculture in GNP tend to decrease

Oct 2008 Oct 2008 Jul 2011 Jul 2011 Questions

	Sampling Urban stratification Studied area and selected villages (Vallee, 2008) Subjects and methods Food consumption Anthropometry Contribution of food groups to energy 300 Contribution of food groups to protein Urbanization and intakes content content
	Methods • Spatial disparities of Vientiane ( 189 villages, 4 urban and 3 periurban districts) • Determination of "urbanization level" after 13 variables representing: density, equipment, accessibility; housing; activity… from the 1995 and 2005 census (Vallee, 2008) • Selection of 148 urban villages (Hierarchical classification analyze) • Determination of 3 urban spaces (Correspondence analysis with a hierarchical forward design, Vallee, 2008) Division of 148 villages into 3 strata: • Central zone: 25 villages • First belt of urbanization: 67 villages • Second belt: 56 villages • Sampling: youngest , pregnant or lactating women, from the whole sample: 157 mother-child pairs • Repeated visits: 2007, 2008 and 2009; home to home surveys • Food consumption survey (24 hours recall and frequency questionnaire) • Anthropometry and clinical examination • Use of Nutrisurvey software (www.nutrisurvey.de) • Compilation of data fromThaï, USDA, FAO, Canadian and German data base • Survey file: 252 entries. • More than 100 nutrients practice analysis of 23 nutrients and energy • Food consumption scores 0 0 scapul, iliac) 0 10 skinfolds (tric, bic, sub 50 20 circumference, 4 10 could be analyzed. In • Children: Weight, length (supine), head and arm circumferences, triceps skinfold • Mother: weight, standing and sitting height, arm Results t 30 100 150 e r c 20 n t e 30 c e 40 diet P e r 50 • But differences in the structure of the 40 200 60 P urbanization 50 70 250 • No quantitative differences according to 60 80
	Energy PUFA Vit A Vit B1 Vit B2 Vit B6 Folates VitC Meat Fish Cereals Vegetables Legumes Eggs Fruits	Ca	Fe Milk	Zn
				19 20

• Are signs of a nutritional transition detectable in food eating pattern and nutritional status? • Are they related to urbanization process?

Conclusions