Welding of a metal-polymer laminate

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Welding of a

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The research described in this thesis was carried out in the framework of the Strategic Research Programme of the Netherlands Institute for Metals Research in the Netherlands (www.nimr.nl).

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Welding of a

Metal-Polymer Laminate

PROEFSCHRIFT

ter verkrijging van de graad van doctor aan de Technische Universiteit Delft,

op gezag van de Rector Magnificus prof.dr.ir. J.T. Fokkema, voorzitter van het College voor Promoties,

in het openbaar te verdedigen op 19 december 2007 om 10.00 uur

door

Heather Laurie GOWER

Master of Applied Science

Ballistic Impact Response of Woven Kevlar Composites University of Waterloo, Canada

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iv

Dit proefschrift is goedgekeurd door de promotor: Prof.dr. I.M. Richardson

Samenstelling promotiecommissie: Rector Magnificus, voorzitter

Prof.dr. I.M. Richardson, Technische Universiteit Delft, promotor Prof.dr.ir R. Benedictus, Technische Universiteit Delft

Prof.dr.ir J. Meijer, Universiteit Twente

Prof.dr.ir A. Verkooijen, Technische Universiteit Delft Prof.dr H. de Wit, Technische Universiteit Delft Dr.ir T. van der Veldt, Corus RD&T

Dr.ir R.R.G.M. Pieters, Technische Universiteit Delft

Published and distributed by: Ponsen & Looijen b.v. P.O. Box 68 6700 AB Wageningen The Netherlands Telephone: +31 317 423 107 E-mail: wageningen@p-l.nl ISBN 978-90-77172-346

Keywords: Pulsed laser, Nd:YAG laser, Welding, Metal-polymer laminate, Sandwich structure

Copyright c 2007 by H.L.Gower

All rights reserved. No part of the material protected by this copyright no-tice may be reproduced or utilised in any form or by any means, electronic or mechanical, including photocopying, recording or by any information stor-age and retrieval system, without written permission from the publisher: Ponsen & Looijen b.v.

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Introduction

Since ARALL (Aramid Reinforced ALuminum Laminates) and GLARE (GLAss REinforced aluminium laminate) were developed in the early 1980s, metal-polymer laminates have found a niche in the aerospace, automotive, and medical sectors where the combination of lightweight, high strength, and stiffness is desirable. In addition to ARALL and GLARE, metal-polymer sandwich materials such as Hylite and Steelite have been developed. Unlike the metal-acrylic laminates used for years to reduce vibration in vehicles, these laminates have polymer cores which are up to five times as thick as the metal layers.

The metal-polymer laminates are generally made up of sheets of metal, either aluminium or steel, between 0.1 and 0.2 mm thick, layered with a polymer layer. The polymer layer is often polypropylene or polymer com-posite such as high strength glass or Kevlar fibre in an epoxy matrix. The polymer layer is generally about 0.8 mm thick.

In addition to all of these materials having low weight and good physical properties, they also share another characteristic; they are difficult to join. The difficulty with joining the laminates either to themselves or to other materials is one of the main obstacles to these laminates being widely ac-cepted. Prior to the work presented in this thesis, the only joining options were mechanical joining, such as riveting, adhesive bonding, or a welding process that destroys the polymer entirely in the weld region.

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2 Chapter 1. Introduction

a sandwich structure, made up of two steel sheets as the skin layers and a polypropylene core is considered. Steelite is a prototype material produced by Corus with 0.12 mm thick mild steel skin layers, and a 0.62 mm thick polypropylene core.

Due to the thickness of the steel sheets and the low degradation tem-peratures of the polypropylene, a low heat input process such as pulsed Nd:YAG laser welding is required for welding of the skin layers. Use of this process coupled with scanning optics makes it possible to control not only the amount of power throughout each pulse, but also the location of each weld. The stitching method takes advantage of weld placement in such a way as to maximise cooling time between pulses and thus minimise heat build-up.

Although pulsed laser welding is a low heat input method, some con-cerns still remain. Pulsed laser welding results in high cooling rates which could lead to problems with cracking and the formation of unfavourable mi-crostructures. Also, it is unlikely that degradation of the core layer can be eliminated completely, so there will be hydrocarbon degradation gases, these gases may negatively influence the weld quality in terms of porosity and weld composition. Additionally, although each spot weld has a low energy input, the overlapping of these welds could lead to a high heat input and heat build up, increasing the potential for core degradation. Further, the mechanical properties of full penetration welds with increased core degradation, as com-pared to partial penetration welds with minimal degradation must also be considered.

This thesis is comprised of several parts. The following chapter gives an overview of the background information required to support the rest of the work. Very little information directly relating to the welding of metal-polymer laminates could be found, therefore the supporting information focuses on a range of related aspects including: material properties; the influence of process variables; weld defects; fluid flow; and modelling of conduction mode welding.

The third chapter gives a description of the equipment used during the experimental work, and also describes metallographic sample preparation, and various test methods used to analyse the welds.

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3

of high speed video experiments, tensile tests, and hardness tests are also presented.

The results from the fourth chapter are discussed in chapter five in terms of weld dimensions, influence of each welding method on the polymer core, and resultant microstructure. Additionally, the influence of welding pa-rameters on the final weld dimensions and properties is addressed and a comparison of the three welding methods is made.

The sixth chapter focusses on how welding of the skin layer affects the polymer core. It order to enhance the understanding of the polymer core, the results of differential scanning calorimetry and thermogravimetric analysis are presented. The damage to the core layer is discussed and related to welding parameters and welding method. Defects in the weld metal directly resultant from the core degradation are also discussed.

Chapter seven presents the results of analytical and thermal finite ele-ment models of pulsed laser welds. Linear regression relationships are de-rived to show which welding parameters have the most significant influence on resultant weld dimensions. The arguments for and against each type of model are discussed in terms of accuracy, ease of use, and versatility.

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Chapter 2

Background

2.1 Metal-Polymer Laminates

In an effort to develop light weight high stiffness materials for the aero-space industry, Schijve and Vogelesang at Delft University of Technology de-veloped two fibre metal laminates, ARALL (Aramid Reinforced ALuminium Laminate) and GLARE (GLAss REinforced aluminium laminate) [1]. These materials were some of the first Metal-Polymer Laminates (MPLs) available commercially, being produced by the Structural Laminates Company at the beginning of the nineteen nineties [2]. Both ARALL and GLARE are mul-tilayer laminates, with alternating layers of 0.2 to 0.5 mm thick aluminium sheet and a fibre composite. In the case of ARALL an aramid/epoxy com-posite is employed, and an S2 or R-glass/epoxy comcom-posite in the case of GLARE [2, 3].

In 1994 the sandwich material, Hylite was added to the market by Hoogovens Hylite BV, an affiliate of Koninklijke Hoogovens [4]. One of the first applications of Hylite was in the automotive industry as the roof panel for the Opel Mexx [5]. Hylite is an MPL with 0.2 mm thick AA 5182 skin sheets and a 0.8 mm thick polypropylene core [6]. Steelite, the steel equivalent to Hylite with 0.12 mm thick mild steel skin sheets and a 0.62 mm thick polypropylene core has also been developed; however this material is not yet commercially available [5].

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6 Chapter 2. Background

These laminates have different specific advantages depending on the ma-terials, although they all share general advantages of high stiffness and light weight.

Joining methods for metal polymer laminates have been limited to me-chanical joining, such as riveting, adhesive bonding, and splicing [1, 6]. All of these methods increase the thickness at the joint as they require an overlap. The literature reveals no method for butt welding metal polymer laminates where the skin layer is a fraction of the thickness of the core. Laminates made up of a thick skin and a thin polymer core are sometimes used for their vibration damping characteristics, these laminates have been success-fully welded by means of resistance spot welding and friction stir welding, however both of these methods are limited by the thickness of the core [7, 8].

2.2 Material properties of Steelite

The laminate studied in this work is Steelite, which is made up of 0.12 mm thick mild steel skin layers and a 0.62 mm thick polypropylene core. An understanding of the thermal and mechanical properties of these component materials is necessary for experimental design and for numerical modelling. For this reason the material property data, as determined from the literature, is presented here for both materials.

The steel used in Steelite is a low carbon, low alloy steel; unfortunately the grade of the steel is not known. For the purposes of numerical analysis the material properties of low carbon steels and iron will be used, these are shown in Table 2.1.

The core material of Steelite is polypropylene, a thermoplastic polymer. Polypropylene was chosen for its relatively high temperature resistance as compared to other thermoplastics [4]. By using polypropylene (PP) it could be guaranteed that the Steelite would remain stable at 423 K for at least 30 minutes allowing for paint hardening [6].

Propylene is a member of the vinyl group, which means that one of the hydrogen atoms in ethylene is replaced with a side group, in this case CH3. When the double bond is broken and the monomers form a chain, or

a polymer, polypropylene is formed [10]. Figure 2.1 shows the structure of polypropylene.

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Table 2.1: The thermal properties of AISI 1006 [9]

Temperature Thermal Conductivity Specific Heat Emissivity Density

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Figure 2.1: Structure of polypropylene [10].

Table 2.2: Typical properties of polypropylene [11].

Material Property Value

Tensile Strength [MPa] 40

% Elongation 700

Elastic Modulus [GPa] 1.46 Density [kg·m−3

] 900

Izod Impact [J·m−1

] 53

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Table 2.3: Thermal conductivity, density, and heat capacity of polypropylene ho-mopolymer [12]

T Thermal Thermal Volumetric

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The temperature at which polypropylene melts or begins to degrade is affected by the heating rate. At low heating rates the melting temperature may actually be higher as crystallisation occurs during heating. For example at a heating rate of 10 K·min−1

the melting temperature is 428.0 K, whereas at 100 K·min−1

the melting temperature is 422.3 K [13]. The melting and degradation temperature is dependant on the structure of the polypropylene, and melting temperatures of up to 440-450 K have been reported [14, 15, 16]. Degradation temperature tends to increase with heating rate, for example, at a rate of 5 K·min−1

the degradation temperature, in an argon atmosphere, is 714 K, however, at a rate of 20 K·min−1

, the degradation temperature is 743 K [17].

The products of degradation must also be considered. If the polypro-pylene ignites, which is unlikely if properly shielded from oxygen, the degra-dation products would consist mainly of CO2, H2O and CO [18].

Degra-dation in an inert atmosphere is more likely. In this case the degraDegra-dation products were found by Guddeti et al. [19] through the use of an induction-coupled plasma (ICP) reactor, to consist of 94% propylene, 2.5% methane, and 2% ethylene, the balance was made up of various C4s such as butane. The ICP reactor uses high temperatures (3000-8000 K) and high heating rates (106 _K·s−1

) and thus is more likely to predict the degradation prod-ucts that would be seen during welding. Other studies, which involve lower temperatures and longer times have shown that reaction products include branched methyl hydrocarbons [20], and other unsaturated low-molecular weight hydrocarbons [21].

For laser welding latent heat is also of interest. This can be determined through differential scanning calorimetry (DSC). For fully crystalline poly-propylene, the latent heat of melting is 209 kJ·kg−1

[22].

2.3 Laser spot seam welding

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Figure 2.2: The surface of a pulsed Nd:YAG weld [25].

2.3.1 Pulsed laser spot welding modes

Pulsed laser welding can occur in two modes, the first is conduction mode welding, where the boiling temperature of the base material is not reached, the second is a keyhole mode where the base metal is vaporised and a column of vapour is generated inside the melted area. The efficiency of keyhole mode welding is much greater than that of the conduction mode due to multiple reflections of the laser beam inside the keyhole [26]. For aluminium, the melting efficiency is less than 1% in conduction mode, however, in the keyhole mode it is about 10% [27]. For steel, reported efficiencies vary from 2% to 67% depending on process characteristics such as welding mode, surface finish, and laser power, among others [27, 28, 23, 29, 30].

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is that keyhole spot welds tend to have porosity [30]. The porosity is caused by the keyhole collapsing and trapping gas at the base of the weld pool [31]. Keyhole mode welds also tend to have craters due to loss of material from vaporisation [32]. Unlike keyhole mode welds, conduction mode welds tend to have fewer defects [32], and are expected to be of greater relevance for the present study.

2.3.2 Process variables and their effects

There are several process parameters that can influence the quality of laser welds, such as shielding gas, ambient temperature, heat source param-eters, etc. Only the most important ones will be discussed in this review. These include: source power and intensity; the effect of focussing; pulse duration; and pulse shaping.

Power and intensity

The peak pulse power and intensity are two of the most important pa-rameters in pulsed welding. The intensity refers to the power per unit area, so a smaller spot size gives a higher intensity then a larger spot size for the same power. A higher intensity also leads to an increased energy density for the same pulse duration.

Higher intensity pulses create spot welds with deeper penetration and, in some cases, larger diameters [33]. Figure 2.3 shows the effect of power intensity on the penetration during pulsed welding of 304 stainless steel. Data for two different pulse times are given.

Manipulating the power can help to achieve desired results, for example using high peak powers and low pulse times can reduce peak temperatures [33] while still creating good quality welds. This is important when welding temperature sensitive components, such as pace maker batteries where high temperatures can destroy glass to metal seals [33].

The power distribution is also important. Models show that final weld dimensions are very sensitive to energy distribution [9]. A model using a Gaussian heat source will more accurately predict weld dimensions than one using a top-hat profile. Tzeng determined that it is the average peak power density, given by Equation 2.1, that affects the final weld shape the most [34].

Average P eak P ower Density=

T otal P ulse P ower (J) P ulse Duration (ms)

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2.3 Laser spot seam welding 13 0.0 0.5 1.0 1.5 0 500 1000 1500 2000 2500 Power intensity (Wǜm-2 ) P e n e tr a ti o n d e p th ( m m ) 7.0 ms 2.2 ms

Figure 2.3: Effect of power intensity on penetration depth of pulsed Nd:YAG welds with two different pulse times on 304 stainless steel [23].

Focusing

Focusing conditions are determined by the focal distance, which is de-pendent on the design of the lens and the location of the focal point.

The focal distance affects the weld width and also the effect of other parameters. Longer focal distances result in larger focal spot diameters [35], producing larger weld widths [33].

It is possible to weld with the focal point above, below, or at the work-piece surface (Figure 2.4). The technique of welding with the focal point above the surface is called positive defocusing, whereas when the focal point is below the surface it is considered to be negative defocusing. The location of the focal point not only determines the size of the spot, and therefore the power intensity, but also determines whether the laser beam is divergent or convergent at the surface. For example if positive defocusing is used, the laser beam is divergent at the surface of the material. This means that with increasing depth the intensity decreases. This is beneficial when trying to achieve a stable keyhole or make shallow, wide welds [31]. Wider welds can be beneficial as they reduce the sensitivity to fit-up.

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POSITIVE

NEGATIVE

focal point

a

a=a

Figure 2.4: Positive and negative defocusing.

instability and gas entrapment [30].

The size of the focal point has a significant effect on the weld bead size. In general, the weld pool width is 50%-100% greater than the focused spot diameter [23]. Fuerschbach and Hinkley found that the use of different focal lengths affected how much a variation in other parameters, such as pulse power and pulse duration, would change the depth of penetration. They indicated that the use of a lens with a 160 mm focal length, as opposed to a 120 mm or 200 mm focal length, reduced the variation in depth of penetration when other parameters were varied [33]. A longer focal length results in a larger spot size and depth of focus, and since the depth of focus increases from the 120 mm lens to the 160 mm lens, it is not surprising that the variation in weld depth decreases, as slight variations in distance between the lens and the workpiece will have less of an effect on depth of penetration. The reason that the 200 mm lens resulted in more variation in the depth of penetration was due to an increase in focal spot size, which was large enough to require an increase in pulse energy.

Pulse time

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For the same net energy, shorter pulse times give deeper penetration [23], this is due to the amount of energy per unit area required to achieve a keyhole. For the same intensity, long pulses give deeper penetration than short pulses [23]. A long, low energy pulse will generally create a conduction mode weld. With all other parameters such as pulse power and energy density held constant, penetration, aspect ratio, and diameter all increase with pulse time [32].

Liu et al. observed that no further increase in weld size occured with pulse times over 4 ms for 3 mm thick AISI 409 steel, which means that there is no benefit to using pulse times above 4 ms for either conduction or keyhole mode welding for this material thickness and range of power densities [32]. This study also found that the rate of metal loss decreased after the first 3 ms of a pulse [32]. When investigating AA 1100 aluminium, Weckman et al. [27] found that conduction mode welds showed a rapid increase in size with pulse durations up to 2 ms, after which there was no further increase.

The pulse time was also seen to affect the quality of the weld. Figure 2.5 shows how pulse time influences the amount of porosity and crater formation present in a weld. It can be seen that for low pulse times there is no porosity, but a large crater is formed.

0.0 0.5 1.0 1.5 2.0 0.5 1 1.5 2 4 8 12 16 Pulse time (ms) W e ld , c ra te r, a n d p o ro s it y a re a ( m m 2) Weld Porosity Crater

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Pulse shaping

The temporal pulse shape can have a profound impact on the quality of a weld. The pulse shape refers to the power versus time relationship over the duration of a pulse. The effect of pulse shaping differs for keyhole and conduction mode welds, pulse shaping has a much more significant effect on the quality of keyhole mode welds than on conduction mode welds [36]. Despite this, only the effect on conduction mode spot welds will be discussed here as keyhole mode welding is beyond the scope of this review, as previously mentioned.

The main reason for using pulse shaping is to reduce defects such as porosity and solidification cracking (which are discussed in Section 2.7). There are several options when choosing pulse shapes, the standard choices are: the normal rectangular pulse; a ramped up shape, where power increases with time; and a ramped down shape where power decreases with time. Several studies have concluded that using a ramped down pulse is optimal for reducing weld defects such as porosity and solidification cracking [37, 38, 39, 36], whereas a ramped up shape promotes higher aspect ratios (weld depth divided by weld width) [37].

The efficiency of pulse shaping depends on the material. For example, a 1994 study on AISI 304 stainless steel concluded that pulse shaping has no effect on conduction mode spot welds [36]. Another study discovered that pulse shaping did affect the quality of pulse welds on STS 310S stainless steel [37], however, is was not clear if the welding mode was conduction or keyhole. The literature indicates that pulse shaping always affects the quality of aluminium pulse welds, especially in alloys susceptible to cracking [38, 39]. The optimised pulse shape for a solidification crack susceptible aluminium alloy, as found by Michaud et al. for a 5 mm thick sample is shown in Figure 2.6. The steps are a result of the optimisation method which divided the pulse into 20 equal sections.

2.4 Microstructure

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2.4 Microstructure 17 0 500 1000 1500 2000 0 5 10 15 20 25 Time (ms) P e a k p o w e r (W )

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thin band of planar growth at the fusion boundary where the cooling rate is the lowest [38]. When there is slight undercooling, solidification proceeds in a cellular manner in the direction of the heat gradient. As the cooling rate is increased, the solidification occurs by columnar dendritic growth, then by growth of equiaxed dendrites, these are also shown in Figure 2.7.

Figure 2.7: Effect of cooling rate on solidification microstructure [40].

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In pulsed laser welding, the cooling rate is very high, which tends to result in a fine microstructure. Uenishi et al. [41] found that in pulsed Nd:YAG welding of 0.25 mm diameter austenitic stainless steel wires, the microstructure was columnar dendritic, with the arm spacing suggesting cooling rates of 2x105 to 2x106 _K·s−1

. Michaud et al. [38] found that, with a rectangular pulse shape applied to an Al 3.73 wt.% Cu alloy, there was a 3-5 µm thick cellular dendritic band followed by a very fine microstructure, with a cell spacing of less than 1 µm. When a ramped down pulse shape was used the microstructure changed to a 5-10 µm thick planar band, followed by a 50-70 µm thick band of cellular dendritic growth, with the centre of the weld being a finer microstructure. This shows that the pulse shape significantly affected the cooling rate.

Fine microstructures are generally considered beneficial, improving duc-tility and yield strength [40], however there are also disadvantages to high so-lidification rates. One of these relates to soso-lidification cracking. If the metal solidifies too quickly the molten metal cannot flow to heal any voids caused by too fast dendrite growth, this is why the fine microstructure resulting from a rectangular pulse shape, as opposed to the coarser microstructure from a ramped down pulse shape, can lead to solidification cracking [38]. Additional problems associated with high solidification rates are the forma-tion of brittle martensite in steels, which can act as a crack initiator, and segregation of the alloy during solidification [10].

2.5 Fluid flow in weld pools

There are three main driving forces for convective fluid flow during weld-ing where a keyhole is not present, these are: electromagnetic (Lorentz) caused by the interaction of the divergent current path and the magnetic field generated by the current flow; buoyancy caused by temperature gradi-ents in the weld pool; and surface tension gradient (Marangoni) caused by temperature variation at the free surface and influenced by the composition of the material being welded [42, 43].

Electromagnetic force has a strong effect on weld pool shape as a fluid flow loop which goes down in centre of the weld pool and up at the sides is created, favouring deep penetration and narrow welds [42]. However, this does not apply to laser welding since there is no current flow [43].

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the weld pool the liquid metal cools, and thus becomes more dense causing it to sink. The liquid rising in the middle and sinking at the edges creates a flow loop in the weld pool, as shown in Figure 2.8, favouring a wide shallow weld pool. Buoyancy is reported to be one of the weakest of the driving forces for fluid flow in weld pools [42].

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2.6 Oscillation in weld pools

During welding the molten weld pool oscillates, the cause of the oscilla-tion is the variable pressure from the arc or the laser. In a pulsed process the application and release of this pressure initiates oscillation by depressing the liquid, and then by the recoil of the liquid when the pressure is released. Damping of these oscillations occurs during solidification, as does dissipation of the energy due to the viscosity of the liquid [45].

There are several modes of oscillation; those most commonly observed during spot welding are modes one, two, and three. Mode one oscillation, shown in Figure 2.9, is an up and down movement of the weld pool and is seen mostly when the arc/beam and the work piece are stationary and the weld pool is round. Mode two oscillation, shown in Figure 2.10 is a back and forth movement of the liquid, this is generally observed when either the arc/beam or the workpiece is moving, causing the application of pressure to be off-centre with respect to the pool axis [45]. Mode three oscillation refers to full penetration welds oscillation in an up and down movement. Xiao de-rived simplified forms of the equations predicting oscillation frequency (f ) for these modes in both partial and full penetration weld pools (Equations 2.2 to 2.6 [45]).

For partially penetrating welds: mode 1:f2 = 1 4π2 11.04 deq g + 1345.5 d3 eq γ ρ tanh 22.08H deq (2.2) mode 2:f2 = 1 4π2 7.66 deq g +449 d3 eq γ ρ tanh 15.32H deq (2.3) Which can be simplified as:

mode 1:f = 5.84 γ ρ 12 d−32 eq (2.4) mode 2:f = 3.37 γ ρ 12 d−32 eq (2.5)

For fully penetrating welds:

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2.6 Oscillation in weld pools 23

where g is the gravitational constant, deq is equivalent diameter, H is

weld pool depth, γ is surface tension and ρ is the density of the liquid. Equations 2.2 to 2.6 are valid given the following assumptions: the liquid is incompressible, the fluid flow is irrotational, the vertical motion is small compared to the oscillation wavelength, and the weldpool diameter is less than 10 mm in the case of Equations 2.4 and 2.5.

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2.7 Welding defects 25

2.7 Welding defects

There are several types of defects common during welding, these include porosity, crater formation, cracking, and in the case of laminates, delami-nation. This section discusses some of the causes and method of prevention for defect initiation and propagation.

2.7.1 Porosity and crater formation

Porosity is a concern because pores reduce the area of the weld, create stress concentrations, and are crack initiation sites. Irregular pores are a more serious concern then spherical pores, as they lead to higher stress concentrations.

Porosity can be a significant problem during pulsed welding. The two main types of porosity found in pulsed welding are large occluded pores at the root of a keyhole weld, shown in Figure 2.11, or in the centre of a weld, and micro pores along the fusion boundary, shown in Figure 2.12.

One source of porosity is rejection of dissolved gases by the solidifying weld metal. For example, at 1500 K solid iron has a maximum nitrogen solubility of 4 wt.% and an oxygen soluability of 0 wt.%, whereas liquid iron has a nitrogen solubility of 8 wt.% and above, and an oxygen soluability of 23 wt.%, this means that at least 4 wt.% nitrogen and up to 23 wt.% oxygen is rejected on solidification [46]. At high solidification rates the rejected gases can be trapped as porosities as there is insufficient time for te gases to escaple through the molten weld pool[40].

The welding mode and presence of dissolved gases are not the only fac-tors that affect pore formation. The frequency of occurrence of craters and porosity was found to increase with increased power intensity due to an increase in vaporisation and melt expulsion [32], as shown in Figure 2.13.

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Figure 2.11: An example of porosity caused by an unstable keyhole [31].

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2.7 Welding defects 27 0.00 0.05 0.10 0.15 0 2 4 6 8 10 12 14 Power intensity (GWǜm-2₎ P o ro s it y a re a ( m m 2) 4 ms 8 ms 12 ms 16 ms

Figure 2.13: The affect of power intensity on the amount of porosity in AISI 409 laser spot welds [32].

Crater formation occurs when a weld is terminated improperly. Crater-ing occurs due to loss of metal from vaporisation, and contraction of the weld metal during cooling. Fuerschbach and Eisler found that a mass of ap-proximately 165 µg is lost per pulse [23] during laser welding of 304 stainless steel for pulse duration of 2.2 to 7.0 ms and average pulse powers of 40 to 110 W. The resultant crater size was not noted.

2.7.2 Cracking

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There are several ways in which solidification cracking may be controlled. The first of these is by using a preheat. Preheating reduces the cooling rate, allowing voids to seal prior to solidification [48, 47]. Preheating is not an option when considering metal-polymer laminates, however, due to the risk of melting and degradation of the core layer.

The second method is through control of impurities. In steel the po-tential for solidification cracking is increased by the presence of sulfur and phosphorus as these elements are rejected into the liquid during solidifica-tion and form low melting temperature sulphide or phosphide films along the interdendritic regions [49]. The presence of manganese tends inhibiting the effect of sulfur by forming higher-melting temperature manganese sul-fide eutectics in preference to FeS, reducing solidification cracking [49, 47]. Solidification cracking can thus be reduced through composition control and limitation of impurities.

The third way to control solidification cracking applies specifically to pulsed welding. By using pulse shaping the cooling rate, and thus the mi-crostructure of the weld, can be controlled. A ramped down pulse shape has been proven to be able to eliminate solidification cracking completely in both Al-Cu alloys, and 5000 series aluminium alloys [38, 50]. Figure 2.14 shows the improvement in quality in a pulsed weld from pulse shaping.

Figure 2.14: Effect of pulse shaping on weld quality. The weld on the right was made using a rectangular pulse, the one on the left was made with a ramped down pulse shape [38].

2.7.3 Delamination

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skin layer, decreasing the stiffness of the material [51]. Unfortunately, due to the novelty of the materials in question, no literature is available on thermal delamination in metal-polymer sandwich structures.

2.8 Modelling of Laser Welding

Modelling is gaining importance in many areas due to its potential for cost reductions. A good model can predict the outcome of experiments without requiring the consumption of large amounts of material and time. There are several types of modelling that can be used to help predict results in real life situations including physical modelling, where a scale model of the object or process in question is built, or mathematical modelling. Many types of mathematical models are available, such as analytical models and finite element or finite difference models. Neural network models and sta-tistical models are also possible options, but will not be discussed here as they both require significant amounts of experimental data [52].

Physical models (eg. a scale model) are useful although expensive. A physical model can confirm numerical predictions or can be used when math-ematical predictions are difficult due to the number of variables involved. Physical models, however, have limited use for modelling welding as scaling down is impossible in some cases, and can change results dramatically in others.

It can be expensive to conduct experiments to optimise a process. By using mathematical modelling in conjunction with experiments, the number of experiments and the cost can be greatly reduced.

Analytical models consist of a system of equations that has a solution. These models are elegant and can provide good approximations, however they tend to rely on assumptions to ensure their simplicity. For example, in welding, analytical models tend to ignore such factors as fluid flow, ac-tual heat source distribution, latent heat, or the temperature dependence of physical properties.

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depend upon significant advances in processor speed and memory capacity in computers. Most models now focus on one of the significant areas such as fluid flow, microstructure evolution, or keyhole formation (for laser welding). In this work thermal modelling will be used to assess the potential for degradation in the core material under conduction mode laser spot welding in laminate materials. A multi layer system is complex and the bound-ary conditions required to accurately model such a system come into direct conflict with the simplicity assumptions required for the application of the majority of analytical models, thus this section will focus on numerical mod-els; however, an overview of some analytical models will also be given.

2.8.1 History of modelling in laser welding

One of the earliest models used for prediction in welding is an analytical heat flow model developed by Rosenthal in the 1940’s [9, 53, 54]. Rosenthal’s model requires several assumptions, these are:

• A moving heat source on the plate surface • Quasi-steady state

• A source travelling at a constant velocity

• Heat is input as a point (3D solution) or distributed over a line (2D solution)

• The physical properties are temperature independent

• Cooling occurs only by conduction, that is to say, radiative and con-vective heat losses are neglected

• No phase changes occur

Rosenthal accounted for both the two dimensional case of a thin plate, and the three dimensional case of a plate of semi-infinite thickness. In the two dimensional case the heat source is modelled as a line and the temperature profile can be calculated according to Equation 2.7.

T − T0= Q 2πkSd exp −xV 2αS K0(φ) (2.7)

where T is the temperature, T0 is the initial plate temperature, Q = η ·P

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plate, the subscript S indicates the solid phase, αS is the thermal diffusivity

of the plate, d is the plate thickness, V is the welding speed, φ = _2αrV

S, r is the radial distance from the arc to the point at which T is located, and K0(φ)

is the modified Bessel function of the 2nd kind, order zero with argument φ. The three dimensional case, which models the heat sources as a point, is described in Equation 2.8. T − T0 = Q 2πkSRxyz exp −(x + Rxyz)V 2αS (2.8) where Rxyz = (x2+ y2+ z2) 1

2, in which x, y, and z are coordinates with respect to the moving source, which is located at the origin of the coordinate system.

From the assumptions given and the equations, it can be seen that this model has several limitations. It does not account for any microstructural variation, heat flow, latent heat, or convective heat losses. In the case of a thin plate the heat loss due to conduction would be limited and convec-tive heat losses would be dominant, therefore the accuracy is limited. The lack of temperature dependent thermal properties is also a problem. Due to this Rosenthal’s model should only be used below 20% of the melting temperature of the plate in question [55].

Although Rosenthal’s model was originally developed for GTA welding, it can also be applied to laser welding. However, since there is no method to account for fluid flow or vapour pressure, the model cannot predict keyhole formation.

In 1973 the first model specifically for laser welding was developed by Swift-Hook and Gick [56]. This model neglected keyhole formation and fluid flow within the weldpool. The model was improved upon by the addition of a plasma filled keyhole in 1976 by Klemens [57]. Also in 1976 Andrews and Atthey developed a three dimensional model which accounted for convective flow in the weld pool [58]. Although this model incorporated the effect of gravity and surface tension, the affect of buoyancy and the plasma were neglected [9].

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Davis et al. implemented fluid flow in their model in 1986 by assuming that the Peclet number was small, and that the Navier-Stokes equations applied [59].

Mazumder and Steen developed the first numerical model of a continu-ous laser welding process. This model did not consider convective flow, or temperature dependent properties, but did account for the change in ab-sorptivity of incident laser energy with the change from the solid to the liquid phase. An approximation of 20% absorption below the boiling point and 100% above was used. A Gaussian distribution was employed to model the heat source [60].

Further advances include the description of attenuation of laser radiation within a keyhole by Beer-Lambert’s law, and Zacharia’s two-dimensional fi-nite difference model which, although not considering temperature depen-dent properties, accounts for fluid flow (due to surface tension gradient) and calculates thermal cycles and cooling rates [61]. Russo et al. used some temperature dependent properties, such as linearly varying surface tension, and an exponential variation in kinematic viscosity with temperature [9].

In 1999 Frewin and Scott [9] published a three dimensional model which accounts for temperature dependent properties. In addition, an experimen-tally measured laser beam intensity distribution was used, and keyhole for-mation was also taken into account. This model found that the distribution used for modelling the heat source had a significant impact on the accuracy of the results, and that for sufficiently small spot welds, fluid flow could, in fact, be ignored.

2.8.2 Model components

There are some basic components necessary in a finite element model when simulating welding. These include a heat source, defined boundary conditions, and a finite element mesh. This section describes the options for each of these components.

Heat source

Several types of heat sources have been used to model laser welding. Both the physical and temporal distributions of a heat source are of importance. Temporal shaping was discussed in Section 2.3.2, the physical distribution will be discussed here.

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distribution is reasonably accurate for fibre optically delivered laser beams [62].

The most common distribution, is a Gaussian distribution, a high inten-sity in the middle falling off with distance form the origin. The Gaussian distribution is also shown in Figure 2.15.

Distance

Intensity

Gaussian Top Hat

Figure 2.15: Heat source distributions.

A third type of heat source for stationary welds is one which is based on a measured distribution. In this case the actual energy distribution emitted by the laser is measured experimentally and the result is implemented in the model. This option gives the most precise results for a particular laser, but requires remeasurement whenever the configuration changes.

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Boundary conditions and heat loss

Boundary conditions exist in both space and time. An example of a tem-poral boundary condition is the initial temperature of the plate. A boundary in space could simulate clamping through use of rigidly fixed nodes. Bound-ary conditions include mechanical and thermal conditions. In welding, me-chanical boundary conditions are limited to clamping. For spot welds, the effect of clamping, both from a mechanical and from a thermal point of view, is minimal so the focus will be placed on thermal boundary conditions.

There are two types of boundary conditions, Dirichlet and Neumann. A Dirichlet condition proscribes a certain value, such as temperature or position. A Neumann condition does not stipulate a value but rather a slope, such as a heat flux.

In many cases modelling an entire object is not necessary, for example a bead on plate weld is symmetric along the axis of the weld, therefore only half of it need be modelled, reducing computation time. A spot weld is ax-isymmetric, so a two dimensional analysis, rotated around the central axis is sufficient. When symmetry is used, a symmetrical boundary condition must also be implemented. A symmetric boundary condition limits the transla-tion and rotatransla-tion of nodes for mechanical analyses, and can be considered an adiabatic boundary for thermal analyses.

Convection and radiative heat losses can be accounted for by a heat flux boundary condition. This type of condition stipulates that an amount of heat energy, proportional to the difference between the temperatures of the plate and of the surroundings, will be lost. For convection, it is based on Equation 2.9 and for radiation on Equation 2.10. Heat loss due to convection may also be averaged and applied to each node, including the internal nodes.

qc = hAS(TS− T∞) (2.9)

where qc is the convective heat flux, h is the convection coefficient, AS is

the area, in this case of the element, TS is the temperature, of the element,

and T∞ is the temperature of the surroundings.

qr = εσ(TS4− T∞4) (2.10)

where qr is the radiative heat loss, ε is emissivity and σ is the

Stefan-Boltzmann constant.

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temperature. In some cases fixing nodal temperatures can replace mod-elling convection losses by modmod-elling the same losses as conduction losses. This has the potentially simplifying a model.

An example of some of the possible boundary conditions for welding is shown in Figure 2.16.

Symmetric BC - a combination of thermal and mechanical restrictions to simulate a symmetric interface

Neumann BC - heat flux proscribed to simulate convective cooling Dirichlet BC - position proscribed to

simulate clamping Weld

Figure 2.16: An example of some of the possible boundary conditions used during modelling of welding [55].

Meshing

Meshing is very important in finite element analysis. The size of the mesh determines how precisely effects can be calculated, in general the finer the mesh, the more precise the model. However, caution should be exercised when refining the mesh as elements that are too small can produce errors as well.

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36 Chapter 2. Background Symmetry Surface Laser Beam _x y z

Figure 2.17: An example of mesh refining [9].

2.8.3 Models for conduction mode laser spot welding

In this work, only conduction mode spot welds are of interest. Most models for laser welding consider only a moving heat source, or focus on keyhole development. The former is covered by references [39, 30, 63, 64, 65] and the latter by references [26, 66]. This section will focus on existing models for conduction mode laser spot welding.

Models for welding all include a thermal component, all thermal calcu-lations are based on the energy balance equation, an abridged version of which is shown as Equation 2.11 for cartesian coordinates [67].

∂ ∂x k∂T ∂x + ∂ ∂y k∂T ∂y + ∂ ∂z k∂T ∂z + q = ρCp ∂T ∂t (2.11)

where x, y, and z are coordinates, T is temperature, k is thermal conduc-tivity, q is a heat flux, such as convection or radiation, Cp is heat capacity,

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Equation 2.11 can be manipulated for different coordinate systems [68], to include convection or radiation, and to include any other heat gains or losses.

Martin and Bowman

An example of a model for conduction mode laser welding is the Martin and Bowman model developed to predict the temperature field under steady state and transient conditions [69]. By assuming a semi-infinite body, a Gaussian heat source, homogenous material, and by using the heat equation in cylindrical coordinates a temperature field expression was developed.

Equation 2.12 gives the solution for the transient case. It should be noted that by allowing t→ ∞ the steady state solution can be determined.

T (r, z, t) = I0r 2 f 8k Z ∞ 0 J0(ωr)e −r 2 fω 2 8 e−ωz_{erf c} _z 2√αt − ω(αt) 1 2 (2.12) −e−ωzerf c _z 2√αt + ω(αt) 1 2 dω

here I0 is the maximum intensity in the laser beam, rf is the laser beam

radius, J0 is the zeroth order Bessel function, ω ranges from 0 to ∞, r is

the radial distance from the origin, z is the depth, and α is the thermal diffusivity.

This model was improved upon by Weckman et al. [27] by including latent heat and an averaged value for heat capacity as well as incorporating absorptivity. They also used the equation to find a nondimensional absorbed heat flux parameter, given by Equation 2.13.

ξ = 2AbPp πkσg(Tm− T0)

(2.13) where Ab is the absorptivity, Pp is the peak power of the laser pulse, k

is thermal conductivity, σg is a Gaussian distribution coefficient with units

of meters [m], Tm is the melting temperature, and T0 is room temperature.

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Other models

In addition to the Martin and Bowman model, there are several other models for conduction mode laser spot welding available, which will be briefly introduced in this section.

One of the most comprehensive was developed by Chang and Na [70, 71]. This is a three dimensional finite element model for laser spot welding of thin plates that can model both conduction and keyhole mode welding including temperature dependent properties, latent heat effects, keyhole formation, convective and radiative heat loss, as well as some consideration of fluid flow. The material used in the model was stainless steel, and some simplifications, such as assuming the properties of the material above the boiling point were the same as those for iron vapour were made.

Frewin and Scott [9] developed a three dimensional model for a moving pulsed laser. This model used an experimentally measured beam intensity distribution and accounted for the angle of incidence of the source being off axis. Temperature dependent properties and keyhole formation are also accounted for. Latent heat effects were modelled by simulating an artificial increase in the liquid specific heat capacity and emissivity. This model accurately predicted weld depths of both the fusion and the heat affected zones.

Robert and Debroy [72] focused mainly on heat transfer and fluid flow in their model using dimensionless numbers to predict the final weld shape. The dimensionless numbers chosen for the prediction of fluid flow and hence the final weld shape were the Marangoni number, Ma, a measure of sur-face tension driven flow, and the Peclet number, Pe, a measure of forced convection and heat conduction. The thermo-physical properties were as-sumed to be constant for the solid and liquid phases although different values were used for the different states. Seven different materials were welded and modelled to ensure that their results were applicable over a range of physical properties.

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on temperature differences. The model of Katayama et al. is difficult to verify as it is currently impossible to do compositional analyses during laser spot welding, however both the calculated weld depth and weld diameter match the experimentally obtained values within 10%.

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Chapter 3

Equipment and Methods

This chapter describes the equipment and methods used throughout the course of this work. First a description of the laser welding and laser sources is given, this is followed by the procedures adopted for high speed videog-raphy, specimen preparation, microscopy, and various testing methods. De-scriptions of differential scanning calorimetry and thermogravimetric anal-ysis are presented with the experimental results in Chapter 6.

3.1 Laser welding

Throughout the course of this research two different experimental ar-rangements with two different laser sources were used; these are a Trumpf HL204P and a LASAG Easywelder. Laser spot welding was carried out at both locations, however it was only possible to make automated pulsed laser seam welds and stitched welds with the LASAG Easywelder at TU Delft.

3.1.1 Trumpf HL204P - Spot welds

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42 Chapter 3. Equipment and Methods

The beam was inclined at 10◦

from the vertical to avoid reflection of the laser beam back into the lens, and an argon shielding gas with 3% hydrogen at a flow rate of 6 L·min−1

was used. The beam was focussed on the surface of the workpiece using a camera with a focus adjusted so that the focal point would coincide with that of the laser beam.

The workpiece was cleaned with ethanol prior to welding. Non-overlapping spot welds were made, with the weld for each set of parameters being re-peated at least three times. The workpiece was moved after each spot weld ensuring that a spacing of at least 5 mm existed between the spot welds. The time required to move the workpiece and the distance between the welds ensured that the workpiece had sufficient time to cool, avoiding an initial temperature difference.

The facility was equipped with a manually operated x-y table, this facil-itated moving the workpiece precisely and allowed for the manual creation of seam welds for initial tests.

3.1.2 LASAG Nd:YAG pulsed laser - Spots welds, pulsed laser seam welds, and stitched welds

The LASAG Easywelder has a power range of 500 W to 7 kW, maximum pulse duration of 100 ms, maximum pulse repetition rate of 500 Hz, and a maximum pulse energy of 70 J. It also has pulse shaping capabilities with a step size of 0.1 ms. The beam quality is 44 mm·mrad. Instead of a traditional lens system a scanning head is used. The scanning head is equipped with optics with a focal length of 80 mm, a focal spot size of 400 µm, and has a scan area of 30 x 30 mm.

The traverse system has a range of 400 mm on the x-axis, 300 mm on the y-axis, 200 mm on the z-axis. The x, y, and z axes have a repeat accuracy of ±5 µm, a positioning accuracy of ±15 µm and a maximum travel speed of 10 m·min−1

.

Pulsed laser seam welds

The workpiece was cleaned with ethanol, then clamped using a vacuum plate with alignment pins. The experimental setup is shown in Figure 3.1. The laser was focussed using an infrared camera for which the focal point had been adjusted to match that of the laser. The scanning head was then moved to ensure a defocus of -2.75 mm. Argon shielding gas was used for all welds with a flow rate of 12.5 L·min−1

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3.1 Laser welding 43

along the length of the workpiece with travel speeds of 200 or 300 mm·min−1

. All welds were made parallel to the rolling direction.

Figure 3.1: Set-up for pulsed laser seam welding.

Stitched welds

Stitching is a method of creating a pulsed laser seam weld by scanning the laser beam from one point on the surface of the work piece to another. In this way a seam weld is built up of individual spot welds that have been given time to solidify and cool completely before the adjacent pulse is made; hence the degradation of the polymer core is reduced while the overall welding speed is maintained. In addition to the parameters important to the single spot welds, stitching (marking) speed is also a determining factor. One possible ordering of the individual pulses is shown in Figure 3.2.

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Figure 3.2: A comparison of pulse order between a traditional and a stitched pulsed laser seam weld.

between the right side and the middle, so at 0.25 mm and 5.25 mm, then 0.5 mm and 5.5 mm and so forth. The marking speed was set to 10 mm·s−1

.

3.2 High speed video

High speed video was carried out with a Phantom V5 CMOS camera. The acquisition rate was set to 2000 frames per second. The exposure time was 87 µs per frame, and the resolution was 256 × 256 pixels per frame. In order to achieve the magnification required a combination of a Nikon AF-S VR Zoom - Nikkor 70-300 mm zoom lens at maximum extension and a 10x magnification filter was used. The camera was mounted on a tripod and angled at approximately 15◦

from the horizontal. Additional lighting was provided with a white light source with fibre optic delivery from behind the weld. A schematic of the set-up is shown in Figure 3.3. Given the space constraints it was only possible to set up the camera so that it would film one location, while the laser beam travelled from the left to the right of the frame. The camera was manually triggered when welding was started; the time of filming was long enough that a more accurate trigger system was not necessary.

3.3 Specimen preparation

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3.3 Specimen preparation 45 Scanning Head Lens Zoom Lens Workpiece Alignment Pins Highspeed Camera Tripod Arm Vacuum Plate Fibre Optic Light Source

Figure 3.3: Set-up for high speed video.

tests the edges of the specimen were not modified after shearing. After the first series of tests it became evident that the deformation of the edge due to shearing influenced results and the edges were milled from that point on to remove any deformed material.

3.3.1 Grinding and polishing

After welding, the welds of interest were sheared out of the workpiece, these were then mounted in a slow curing epoxy, Struers Epofix/Specifix-20 system. The epoxy is transparent, which allows one to see the surface of the weld, this is of particular use when polishing single spot welds, as it allows for a visual estimation of when enough material has been removed and the centre of the weld has been revealed. To enhance this, the side of the mount was also polished, to give a smooth plane parallel to the surface of the workpiece, in this way a microscope could also be used to examine the surface of the weld during grinding. The second advantage of the slow curing epoxy was that the maximum temperature achieved during curing was well below the melting temperature of the polypropylene.

The single spot weld specimens were ground and polished manually. They were first ground to 4000P with silicon carbide paper. Then the spec-imens were polished to a 1 µm finish with a diamond polishing compound. Before polishing, and between polishing steps, the specimens were cleaned in an ultrasonic bath.

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were first ground with a 200 grit diamond embedded disc and then polished to a 3 µm finish with a diamond polishing compound. A final step was polishing with Struers OP-S, a colloidal silica suspension with a 0.04 µm grain size. The specimens were cleaned in an ultrasonic bath between each step.

3.3.2 Etching

After polishing the specimens were etched. Two different types of etchant were used, the first was picric acid. Specimens were swab etched for 15-25 seconds, and the rinsed and cleaned with alcohol. The second etchant was a 5% nital solution saturated with picric acid. This is a more aggressive etchant. The specimens were swab etched for 10 seconds. Both etchants were used to reveal grain boundaries.

3.4 Microscopy

Prior to sectioning all specimens were photographed using an Olympus SZX9 stereo microscope. The weld width or diameter was also measured. After polishing, but prior to etching welds were photographed with an Olym-pus BX60M optical microscope. Two different light levels were used, one to reveal the metal layer, so it could be ascertained if any porosity was present, and one to reveal the polymer layer. It was necessary to photograph the polymer layer prior to etching to determine the amount of degradation under the weld bead. Measurements for degraded area, depth, and width were taken at this time. After etching the weld cross-sections were again photographed, and the weld width and depth were measured. Welds were photographed as soon as possible after etching. This was necessary because etchant tended to remain in the gap between the polymer and metal where the polymer had degraded, despite careful cleaning and drying after etching, causing over etching of the weld with time.

3.5 Hardness testing

In the Vickers micro hardness test a square pyramidal shaped diamond indentor is pushed into the test material with a given force and at a speed no higher than 0.2 mm·s−1

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3.6 EPMA Line Scans 47

are measured, and the Vickers hardness value for that material is calculated using Equation 3.1.

V ickers hardness = 0.1022F sin

136◦ 2 d2 m ≈ 0.1891 F d2 m (3.1) Where F is the force, dm is the mean diagonal, and 136◦ is the angle

between opposite faces of the pyramidal indentor at the vertex [74].

3.6 EPMA Line Scans

Carbon and manganese line scans were conducted at Corus by means of Electron Probe Micro Analysis (EPMA). The measurements were taken with an SX 100 microprobe from CAMECA. The settings for the carbon analysis were: an accelerating voltage of 5 kV, a beam current of 100 nA, and a beam diameter of 5 µm; for manganese these settings were 15 kV, 100 nA, and 5 µm, respectively. The accuracy of the SX 100 is better than 5% for weight percentages less than 1, this translates into an accuracy of approximately 0.05 wt.%.

Surface contamination accounts for a background carbon signal equiva-lent to 0.6 wt.%.

3.7 Tensile testing

Tensile tests of welds made with several combinations of welding param-eters were carried out. The skin layer of Steelite sheet was butt welded on both sides according to Section 3.1.2. Two tensile test specimens were milled from each sheet so that the weld was perpendicular to the the length of the specimen, and located in the middle of the gauge length. The tensile test specimens were 5 mm wide and 40 mm long, with a gauge length of 20 mm, and are shown in Figure 3.4. In addition to the welded specimens, unwelded specimens were also tested.

In preparation for a tensile test the test specimen was placed in the grips of an Instron 5500R and aligned with the travel axis to avoid incurring any bending stresses. During the test the displacement rate was 2 mm·min−1

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then calculated from the dimensions of the test specimen and the force-displacement curve generated during the test.

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Chapter 4

Welding of the skin layer

-Results

Welding three layers of a metal polymer sandwich material in one step is often impractical due to the different thermomechanical properties of the layers. For this reason a two step approach was chosen, with the first step being butt welding of the skin layer. Three methods for pulsed laser welding of the skin layer were tested, single spot welds, pulsed laser seam welds, and stitched welds. This chapter presents the results for each welding method.

4.1 Single spot welds

A laser spot weld is a precise method for delivering energy to a work-piece, since the pulse duration [ms] and pulse power [W] can be controlled independently, the combination of which results in pulse energy [J]. Pulse shaping, or controlling the power level over the time of the pulse, can also be used, however it was not necessary for the welding of Steelite.

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50 Chapter 4. Welding of the skin layer - Results

Examination of the surfaces and cross-sections of the welds revealed that cracking does not occur for partial penetration welds, as shown in Figure 4.1(a). Once a weld approaches full penetration, a hole through to the polypropylene is formed instead of a weld, a typical example of this is shown in Figure 4.1(b). It is interesting to note that in welds where the penetration is slightly less deep, and a hole has not formed, there is a gap between the polypropylene and the steel under the weld, as shown in Figure 4.1(c). In both these figures a dark circle can be seen, this is a bubble in the epoxy used as a mounting agent, the polypropylene is located below the steel.

The diameter of all single spot welds was measured from the top surface. A selection of welds was chosen for cross sectioning, which showed no ob-vious cracking, and no obob-vious full penetration holes to the polymer layer. After cross-sectioning these welds were polished and etched, and the weld diameters and depths were measured.

The results (Figures 4.2 and 4.3) show that a certain energy threshold must be reached for a weld to be made, and that this energy threshold is larger for larger focal spot sizes. The relationship between weld diameter and pulse energy is shown in Figure 4.2. Each energy value represents a different combination of pulse power and pulse time, thus a different power intensity, which may account for the scatter in the data.

If the threshold for a weld to be made is considered as a threshold energy density, as opposed to a threshold energy input, then there is no longer an explicit dependency on focal spot size. The same data is shown in terms of energy density in Figure 4.3, where a threshold can also be seen, occurring at approximately the same energy density, ∼3 MJ·m−2

, regardless of focal spot size. This indicates that the energy density is the determining factor for the onset of weld formation.

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(a)

(b)

(c)

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52 Chapter 4. Welding of the skin layer - Results 0 100 200 300 400 500 600 700 800 0 0.5 1 1.5 2 2.5 3 3.5 Pulse energy (J) A v e ra g e d w e ld d ia m e te r (ȝ m ) 400 ȝm 600 ȝm 800 ȝm

Figure 4.2: Relationship of pulse energy to top surface weld diameter for three focal spot diameters. 0 100 200 300 400 500 600 700 800 900 0 2 4 6 8 10 12 14 Energy density (MJǜm-2) A v e ra g e d w e ld d ia m e te r (ȝ m ) 400 ȝm 600 ȝm 800 ȝm

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4.1 Single spot welds 53 200 250 300 350 400 450 500 550 600 650 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 Pulse energy (J) W e ld d ia m e te r (µ m ) 0.5 ms 1.0 ms 1.5 ms 2.0 ms 3.0 ms

Welding of a metal-polymer laminate

Welding of a

Welding of a

Metal-Polymer Laminate

Contents

Nomenclature

Constants

Chapter 1

Introduction

Chapter 2

Background

2.1

Metal-Polymer Laminates

2.2

Material properties of Steelite

2.3

Laser spot seam welding

POSITIVE

NEGATIVE

a

a

a=a

2.4

Microstructure

2.5

Fluid flow in weld pools

2.6

Oscillation in weld pools

2.7

Welding defects

2.8

Modelling of Laser Welding

Chapter 3

Equipment and Methods

3.1

Laser welding

3.2

High speed video

3.3

Specimen preparation

3.4

Microscopy

3.5

Hardness testing

3.6

EPMA Line Scans

3.7

Tensile testing

Chapter 4

Welding of the skin layer

-Results

4.1

Single spot welds