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DEDICATION

To my mother, Rama Upadhyay.

iii

ACKNOWLEDGMENTS

Having gone through several drafts, I have realized that it is very difcult if not impossible

to avoid clichs in acknowledgments. Looking back I eventually ended up mostly with

clichs albeit an earnest one. So here it goes.

At the outset I gratefully acknowledge the nancial support of the Center for Friction Stir

Processing which is a National Science Foundation I/UCRC supported by Grant No. EEC-

0437341.

My gratitude is rst due to my advisor Professor Anthony P. Reynolds. This dissertation

would not come to fruition without your constant guidance, support and encouragement.

Thank you, Dr. Reynolds for countless advice and discussion sessions that helped shape

my dissertation and more importantly my understanding of scientic research in general

and science of friction stir welding in particular. I am also grateful for several opportu-

nities including internships, training and conferences you facilitated that has fostered my

professional growth.

I am also indebted to Research Professor Wei Tang and Lab Engineer Daniel Wilhelm.

Dr. Tang has helped me in planning and executing all the welds and provided me practical

guidance on experiment design and testing. No welds would be possible without Dan. Dan

is a master machinist and has been the go to guy for virtually all trouble shooting in the

lab. Thank you, Dan for your time and effort at workshop and out and about.

I also acknowledge Professors J A Khan and T W Knight and Dr. H B Schmidt for their

advice and suggestions while being on my dissertation committee. Thank you all for your

time and effort on going through my research proposal and dissertation. I especially ac-

knowledge thorough and insightful comment on my simulation work by Dr. Schmidt. I also

iv

Figure 4.44 Steady state peak probe temperature vs. tool rpm for welds made

with Densimet and Nimonic Tools.Welds made in 6.35 mm thick

AA7050. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 113

Figure 4.45 Measured steady state torque vs. the tool rotation rate for in air and

under water welds made with Densimet and Nimonic Tools. Welds

made in 6.35 mm thick AA7050. . . . . . . . . . . . . . . . . . . . . . 113

Figure 4.46 Steady state temperature prole from the TPM model along the Z

axis for different tool materials. . . . . . . . . . . . . . . . . . . . . . . 115

Figure 4.47 Temperature eld shown on the transverse cross section of the weld

from TPM model for welds made with different tool materials . . . . . . 116

Figure 4.48 Comparison of experimental tool temperatures for welds made with

Densimet and Nimonic tool . . . . . . . . . . . . . . . . . . . . . . . . 117

Figure 4.49 Vickers hardness proles on transverse cross-section at midplane

for naturally aged samples. Weld performed at 6.8mm/sec using

ceramic tile as backing plate. Forge force used and peak probe T

reached for each cases are indicated. Welds made in 4.2mm thick

AA6056-T451 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118

Figure 4.50 Vickers hardness proles on transverse cross-section at midplane for

naturally aged samples with Aluminum backing plate Welds made

in 4.2mm thick AA6056-T451 . . . . . . . . . . . . . . . . . . . . . . 120

Figure 4.51 Average nugget hardness vs. peak T for welds made over different

indicated backing plates for various welding speeds . . . . . . . . . . . 121

Figure 4.52 Average nugget hardness plotted against peak T for different weld-

ing speeds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 122

Figure 4.53 Change in average nugget hardness after heat treatment plotted against

the peak T. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 122

xvi

welded joints when welded underwater and under sub ambient conditions are presented

and analyzed using various temperature measurements. Since a wide range of peak nugget

temperature were achieved because of the use of various boundary conditions wide win-

dow of weld properties were observed. Grain size correlations among different welds in

through thickness direction are also presented. This has led to a wide extent of correlations

of control parameters like tool rotation speed, welding speed and applied forge force to the

resulting weld properties. Towards the end of the chapter, the concept of composite back-

ing plate is presented. Hardness distributions comparing welds made with conventional

monolithic backing plate and composite backing plate are presented and discussed.

Summary of the results and analysis is presented inChapter 5. Important conclu-

sions and observations that constitute the contributions of this dissertation to the scientic

community is enumerated for all the three alloy avors considered. List of publications

generated from this work is then presented. Challenges ahead in understanding the bound-

ary condition effects in FSW are briey discussed in the end of the chapter. Finally in

Chapter 6 the research approach that might be undertaken in the future to enhance the

understating of the process and its thermal management is presented in some detail. Sev-

eral experiments and simulation works are proposed for the benet of future researchers

interested in pursuing thermal management in friction stir welding.

6

CHAPTER 2

BACKGROUND

2.1 METALLURGY OF 6XXX AND 7XXX A LLOYS

Overview

The fact that certain alloys of aluminum could be made signicantly stronger when rapidly

quenched from a high temperature was accidentally discovered by Alfred Wilm about a

century ago.[15] This opened up a wide area of structural and functional application of

aluminum alloys and made light weight aircraft and locomotive possible. At present a lot

of underlying mechanism that leads to seemingly surprising strengthening is now known

to us. Thanks to the advent and use of advanced techniques such as Transmission Electron

Microscopy (TEM) and Selected Area Diffraction (SAD) that enables high resolution visu-

alization of precipitates, grain boundaries, dislocations and matrices; Differential Scanning

Calorimetry (DSC) that signals phase changes through peaks in heat exchanges at various

temperatures and relatively new Small Angle X-ray Scattering (SAXS) and that provides

quantitative information of precipitate shape, size and distribution, microstructural changes

leading to high strength are now better understood. The eld continues to be an active area

of research and the proposed explanations and results of precipitation sequence are far

from unequivocal, nevertheless certain basic premises are established. It is now known

for instance that the alloying elements like Mg, Zn, Cu, etc. form intermetallic nano and

sub-micron sized precipitates, some of which are primarily responsible for strengthening

by pinning dislocation motion. It has been shown that the distribution, size, shape, vol-

ume fraction and to some extent location of these tiny intermetallic phases are ultimately

7

Recently researchers have considered these alloys as quaternary system and elucidated the

importance of Cu in the alloy system.[22, 23] Using subsequent TEM and SAD analysis

following additional precipitation sequence is also proposed

SSSS GPzones Q(orB;Q) Q(orB;Q)

The metastable needle shaped neb precipitate phase is considered to be responsible for

peak aged condition and indeed has been found to be present in peak strength condition.

[20]In addition to the well knownb phase, a lathe shaped hexagonal precursor phase to

Q phase has also been attributed to strengthening. On the other hand the coarse precipitate

of b which is normally found to form on dispersoids and Q( or B,Q) is thought to be

responsible for overaging as observed by Chakraborty et. al.[22, 21]Similarly AA7050

which is primarily a Al-Mg-Zn-Cu alloy system is has following precipitation sequence:

SSSS GPzones h h(MgZn2)

The major precipitate phase for AA7050-T7 are well established to be stableh( MgZn2)

and/orMg3Zn3Al2 and metastable strengthening phaseh Mg(Zn2;Al;Cu). Other phases

like M(Mg;Zn2;Al;Cu) and S(Al2;Cu;Mg) have also been found to precipitate at high

temperature.[24, 25, 26, 27]h precipitates in addition to some GP zones are established

to be responsible for strengthening. On the other hand coarsening ofh and formation

of non strengtheningh phase that results in depletion of solute from the matrix results in

overaging and strength reduction.2 The precipitation and dissolution sequence at different

temperatures in above families of alloys as they relate to friction stir welding are further

reviewed and dealt in some length later on page 24 and subsequent pages. Table 2.3 below

presents approximate temperature ranges at which corresponding precipitation and disso-

lution events are observed to have occurred. The data are extracted from several research

2For further clarication the reader is referred to sources on precipitation sequencing in age hardeninghardening aluminum alloys.[16, 28]

10

case was 600 mm wide and 600 mm long. See 4.4 on page 149

3.3 WELDING TOOLS

Figure 3.6: Schematic drawings of weld tools used for a) 4.2mm thick AA6056, b) 6.35mm thick AA 7050 and c) 25.4 mm thick AA 6061. Shown by dark dots are the locationsof several thermocouples inside the tools. All Dimensions are in mm

Tools used for production of all the welds were of two piece design with single scroll

typically made out of H13 tool steel and a probe fabricated out of MP-159 (A high tem-

perature cobalt based super alloy) in the shape of a truncated cone (8 taper) with threads

and three ats. For 6.35 mm thick AA 7050 welds, the shoulder was 17.8 mm diame-

ter and probe was 6.1 mm long, with a diameter of 7.9 mm at the intersection with the

shoulder. Same tool was used for 4.2 mm thick AA 6056 by adjusting the probe length

to 4.1mm. For 25.4mm thick AA 6061, a 35mm diameter shoulder was used with 25.2

56

Figure 3.7: Tools used for a) 25.4mm thickAA6061 b) 6.35mm thick AA7050 and 4.2mmthick AA6056.Tool shown in Fig.b has theprobe adjusted for 4.2mm thick welding.

mm long probe. The basic shape and dimensions of the tools are shown in Fig.3.6 with

respective thermocouple placements. The tools used are shown in Fig.3.7.

3.4 TEMPERATUREMEASUREMENTS

Temperature measurements during the welding process were made using K type thermo-

couples attached at various locations of interest. For all the welding performed temperature

was measured by using one or more K type thermocouples (TC) situated at the interior of

the tool. Select welds were also made with plate temperature measurements. Two ther-

mocouple wires (Chromel and Alumel) were welded together using a portable capacitive

discharge welding unit to form a spherical welding bead. The size of the bead obtained was

approximately 1.5mm. Such TC beads were then either spot welded into the tool interior

or glued inside the workpiece at desired locations.

57

Tool temperature data were acquired at a frequency of 1Hz. using a HOBO data logger

(Onset Computer Inc.) A maximum of 4 data loggers can be used simultaneously. The

data loggers were tted on the spindle during welding. Once the weld was complete, the

data was transferred into a computer for further analysis. The thermocouples situated at

the interior of the tool were spot welded into respective holes drilled at desired positions.

The approximate position of TC bead for each tool is shown in the Fig.3.6 by black dots.

Temperature measurements in the plates were made by attaching TC bead at approximate

location of HAZ minimum hardness which typically occurs at about 5-10mm from the weld

seam. Holes were made up to appropriate depth and grooves were made such that the TC

wires were not pinched by the plate under heavy clamping. It has to be conceded that using

thermocouples inside the tool though useful do not provide ideal temperature information.

The uncertainties of using thermocouples in general are listed below.

1. The volume and hence mass of the resulting bead when two thermocouple wires

are welded together can vary and are difcult to control. This may add variation

in the spatial resolution among different thermocouples. The time required to reach

to steady state may depend to some extent on the mass of the bead which is not

controlled. As is the practice beads are only inspected visually for inconsistency in

shape and size.

2. The amount of the glue that goes in to the thermocouple hole for inplate TC is also

not controlled. This may also add up to the uncertainty.

3. There is no direct method to inspect the placement of thermocouples inside the holes.

This may lead to variation in the positioning and orientation of beads among different

installations. Although care is taken during insertion of thermocouple into the hole,

such that the bead reaches similar position, there is no straightforward way to ensure

consistency among different thermocouple placements. The position of each thermo-

couple that is seating on the hole may be different. Given the fact that the gradient in

58

CHAPTER 4

RESULTS AND ANALYSES

4.1 TEMPERATURE, TORQUE AND FORCES

Temperature and torque results from cascaded backing plates

Figure 4.1: Probe temperature trends for two single pass cascaded arrangements of backingplates with a constant forge force of 14.2kN. Backing plate used for each section of theweld is indicated. Bead on plate welds made in 4.2mm thick AA6056-T451

The temperature transients recorded by thermocouple spot welded at the midplane

height in the tool during welds made in 4.2mm thick AA6056 are presented in Fig.4.1.

Welds were made in cascaded backing plate arrangement such that three different back-

ing plate materials are used in the same weld pass while keeping tool rpm, welding speed

and forge force constant. Notice that the time axis starts at 55 seconds because the plunge

sequence is not included in the graph. Although some portion of the probe T does not

62

(a) Al BP. Forge force used: 15.7kN (b) Al BP. Forge force used: 18.5kN

Figure 4.3: Temperature transients measured using thermocouples at tool probe core andshoulder interface for welds done at 320rpm-3.4mm/sec using aluminum backing plate.Welds made in 4.2mm thick AA6056-T451

(a) Ti-6-4 BP. Forge force used: 12.8kN (b) Ti-6-4 BP. Forge force used: 15.7kN

Figure 4.4: Temperature transient measured using thermocouples at tool probe core andshoulder interface for welds done at 320rpm-3.4mm/sec using Ti-6-4 backing plate. Weldsmade in 4.2mm thick AA6056-T451

and 960 rpm - 10.2mm/sec using aluminum and titanium backing plates in Figures 4.3, 4.4,

4.5 and 4.6. These sets represent extreme cases of both welding parameters and thermal

boundary condition at the bottom of the workpiece tested. The peak temperature extracted

from all the graphs like these will be presented later. For the moment few observations can

be made. In general shoulder thermocouple reaches a slightly higher temperature than the

64

(a) Al BP. Forge force used: 12.8kN (b) Al BP. Forge force used: 15.7kN

Figure 4.5: Temperature transient measured using thermocouples at tool probe coreand shoulder interface for welds done at 960 rpm-10.2mm/sec using aluminum backingplate6056-Probe and shoulder T transients

(a) Ti-6-4 BP. Forge force used: 11.4kN (b) Ti-6-4 BP. Forge force used: 14.2kN

Figure 4.6: Temperature transient measured using thermocouples at tool probe core andshoulder interface for welds done at 960 rpm-10.2mm/sec using Ti-6-4 backing plate.Welds made in 4.2mm thick AA6056-T451

probe thermocouple and reaches it faster. At the beginning of the weld corresponding to

the initial phase of the plunge sequence, the probe T is higher until cross over takes place

and shoulder T rises rapidly ahead of the Probe T indicating nal act of plunge sequence

when the shoulder comes in contact with the abutting plates. Some uctuation in shoulder

T can be noted at relatively lower forge forces. This probably is the result of less than ideal

65

One concern with graphs like Fig.4.11 is that it is not only the tool rpm that is varied.

When doing a series of welds by varying tool rotation, FSW practitioner will invariably

have to adjust the applied forge force as well in order to produce quality welds with min-

imal ash. For an empirically established optimum forge force the increase in rpm will

lead to increase in average temperature rendering the material more plastic. Without any

forge force adjustments some of the plasticized material will expel away from under the

shoulder instead of being consolidated behind the tool. This results in excessive ash and

possibly volumetric defect if the forge force is not adjusted as the rpm is increased. The

applied forge force (Fz) may have its contribution in addition to the tool rpm as it directly

affects the frictional heating. It was thought necessary to understand effects of forge force

alone in the process. As discussed in Section 3.2 on page 54, three sets of rpm and welding

speeds were used to make series of welds in 4.2mm thick AA6056-T4. In this series of ex-

periment backing plate material was also varied with the intent to understand the effects of

thermal boundary condition variation at the bottom of the workpiece. In Fig.4.12 the visual

surface quality of welds performed at different forge forces and backing plates are shown

for 640rpm and 6.8mm/sec. Along the X Axis are different backing plates with increasing

diffusivity from left to right ( greater heat loss from workpiece to the backing plate). On the

Y axis from bottom to top are weld surfaces made at increasing forge forces. For a given

backing plate starting with low forge force there is incomplete shoulder contact to begin

with and as the forge force is increased there is full shoulder contact. At the higher end of

the forge force excessive ash is seen. At a given forge force the extent of shoulder contact

and level of ash appears to be dependent upon the heat loss to the bottom of the baking

plate. For instance, whereas weld made with ceramic tile BP and forge force of 12.8kN re-

sults in excessive ash, there is just enough shoulder contact with aluminum backing plate

everything else including the forge force remaining constant. Consequently there is a shift

of good forge force from lower level to higher level of forge force as the change in the

backing plate results in greater heat loss from the workpiece. Cross-section examination

74

Figure 4.12: Representative weld surfaces from series of welds made at 640rpm and6.8mm/sec. Welds made using different backing plates are shown in different columns,while each row represents welds made at an indicated forge force. Welds made with 4.2mmthick AA6056-T4

Figure 4.13: Etched transverse cross-section of welds made using ceramic tile backingplate at 640rpm and 6.8mm/sec using various indicated forge forces.Welds made with4.2mm thick AA6056-T4

75

of these welds show few welds at low forge forces end up with wormhole defects, while

at the high forces excessive thinning due to ash takes place. (See Fig.4.13). The steady

state peak temperature from probe thermocouples for welds in 4.2mm thick AA6056 is

plotted against the corresponding forge force for various backing plates in Fig.4.14. The

trends of temperature is qualitatively similar to the trends shown in Fig.4.11 obtained by

varying tool rpm. Fig.4.14(a) shows the trends of stable peak mid-plane probe temperature

achieved with welds done at 640 rpm and 6.8mm/sec. Excluding bad welds indicated

by cross marks viz: three welds made at relatively low forge forces that resulted in surface

and/or volumetric defects and three welds at relatively high forge force that resulted in high

level of ash and thinning, rest of the runs produced good quality welds with little to no

ash and no defects. This gure illustrates substantial changes in process response that

can be brought about by changes in either forge force, backing plate diffusivity or both

while keeping rpm and welding speed constant. For instance considering only good qual-

ity welds with no defect and minimal ash, combined variation of forge force and backing

plate diffusivity can result in as much as ~70C difference in probe temperature. Change in

forge force or backing plate alone can result in ~50C difference in probe temperature (As

shown by dotted vertical and horizontal lines). At forge axis force of 11.3kN for instance

a ceramic tile BP weld can achieve a peak probe T of 460C, while under equivalent ro-

tational and translation rate steel BP achieves only 431C. Aluminum BP likewise attains

423C. As expected the peak probe temperatures line up in accordance with the backing

plate diffusivity and the applied forge force.

The peak probe T extracted from weld parameter set of 320rpm- 3.4mm/sec and 960-

10.2 mm/sec are plotted in Fig.4.14(b). Indicated are the differences between peak probe

temperatures achieved between two extreme cases of backing plates for two extreme cases

of welding speeds. If the lines of data from different backing plates and same parame-

ter set are assumed to be parallel this shows that the effect of backing plate diffusivity on

achieved peak probe T is greater at lower welding and/or rotational speed and diminishes

76

(a) Thermal history at heat affected zone for weldsmade using different backing plates.

(b) Rate of cooling calculated at the temperaturerange of 320C-250C plotted against the respectivebacking plate diffusivity values.

Figure 4.23: Thermal history at heat affected zone for welds made at 640rpm-6.8mm/secin 4.2mm thick AA6056 for different backing plates.

an appropriate power value,Q, from a typical FSW and the thermal diffusivity of AA7050

at room temperature, temperature elds were obtained for a series of welding speeds. The

times above 200C are plotted with respect to welding speeds in Fig.4.22. The time spent

above 200C is observed to follow a power law relationship with the welding speed thus

illustrating the diminishing returns in the increase in cooling rate with the increase in the

welding speed. Having discussed the effects of ambient conditions on temperature tran-

sients, the effect of backing plate thermal condition on the rate of cooling is considered

next. Temperature transients from the approximate location of HAZ minimum at the plate

midplane thickness obtained from welds made in 4.2mm thick AA6056 at 640 rpm and

6.8mm/s are shown in Fig.4.23a Data are shown for otherwise identical welds made using

ceramic tile, steel and aluminum backing plates. The peak T reached at the TC locations

for all the backing plates are at the range of 355C-380C which is the approximate range

of peak T expected at the HAZ region. The corresponding rate of cooling calculated from

similar temperature transients are plotted against the corresponding backing plate diffusiv-

ities. The cooling rate was calculated for the temperature range of 320-250C which is

86

Figure 4.24: Thermal history at heat affected zone taken at 5 different thickness levelsfor welds made in 25.4mm thick AA6061-T6 using 480rpm-6.8mm/sec. The bottom rightchart shows rate of cooling at different thickness levels calculated at a temperature rangeof 320-250C

relevant for precipitation kinetics in 6XXX series alloys. The graph indicates a gradual

increase in rate of cooling as the backing plate diffusivity increases: ~30C/s for insulating

ceramic tile to 40C/s for highly conductive aluminum backing plate. Similar instrumenta-

tion were also performed on welds made with 25.4mm thick AA6061 at several thickness

levels. Temperature transients measured at approximate locations of minimum hardness at

the HAZ for respective thicknesses are shown in Fig.4.24. The black circles indicate loca-

tions of thermocouples in the welded plate. There is some variation in the peak temperature

reached at each location (320-420C). This large spread in the peak T at approximate HAZ

87

Figure 4.28: Comparison of tensile ow stress vs. temperature among calculated and exper-imental values for 4.2mm thick AA6056-T4. Triangles show tensile ow stress calculatedfrom measured torque plateau while circles show values from Gleeble experiment adaptedfrom [145]

in the torque at both the plunge and regular weld regime. This led both the researchers to

conclude that the contact condition was predominantly sticking rather than sliding. In this

work signicant and consistent change in tool torque with the change in forge force is seen

for each BP as described above. This is in contrast to the results found in the literature

where little to no change is torque was observed with the increase in forge force. This per-

haps means that contact conditions evolved from predominantly sliding to sticking when

the forge force was gradually increased on these particular welds. Considering Fig.4.25,

for all the cases excluding steel and Ti-6-4 backing plate at 960rpm, the torque increases

linearly up to a certain value of forge force suggesting direct dependence of shear force

on the normal force in this regime. The beginning point of plateau region in every case

should indicate the start of the regime where predominantly sticking condition prevails at

the tool workpiece interface. That is the shear stress on the interface is primarily governed

by the workpiece material and not by the tool normal force (Fz ) as is the case for clas-

sical friction. If the ow stress at the tool work piece interface is assumed to be uniform,

torque values can be used to calculate average ow stress of material in contact using the

92

following equation.

t shear=3

2pTshoulderr3o r

3i

Heret shear is the uniform shear stress at the shoulder-work piece interface under the tool,

Tshoulder is the tool torque whilero andr i are outer and inner radii of the shoulder. This

equation assumes uniform shear stress throughout the shoulder interface and torque due to

the probe to be negligible. Equivalent tensile ow stress under the shoulder can then be

calculated by assuming von Misses criteria such thats inter f ace=p

3t shear . At the plateau

region of the torque trend, the calculated tensile stress must be equal to the material ow

stress at a given temperature as per sticking condition.(See Fig.4.27) Since the use of differ-

ent BPs led to different temperatures at which torque plateau occurred: this situation pro-

vides an opportunity to obtain ow stress vs. temperature curve for the workpiece material

specially at a high temperature range. Fig.4.28shows the tensile ow stress values (trian-

gles) calculated from each experimentally observed torque plateau as seen in Fig.4.25. The

data is plotted against corresponding measured peak probe temperature.2 Also shown in

circles are the data set reported by Zain et al. from Gleeble test of AA6056.[145]3 Although

there are only two experimental data points relevant at the temperature range for which

calculated ow stress are shown, there is a good agreement among the two values. As in-

dicated in the chart the experimental data were obtained at the strain rate of 0.002/sec and

0.0002/sec both of which are relatively low values in comparisons to >20/sec anticipated

in FSW. [65] The larger value of ow stress might be indicative of the strain hardening at

greater strain and strain rate operative in FSW. This approach to calculate the ow stress

2Shoulder temperature measurements in 320rpm and 960rpm sets discussed on page 66 show that probeT is a good indicator of shoulder interface temperature for the considered case.

3Gleeble test refers to a wide variety of physical simulations carried out on materials typically at elevatedtemperatures to determine various material properties at different variable parameters. Gleeble systems aim tosimulate thermal and mechanical processes in laboratory that the material is subjected to in actual fabricationor end use[146]. Gleeble test referred above concerns measurement of ow stresses at several elevatedtemperature by varying the strain rates. See [147, 148]

93

Figure 4.29: Measured tool torque transient during last 50mm weld length in25.4mm thick AA6061-T6 for four different backing plates. Parameter setsused for are 480rpm-3.4mm/sec. The average torque is plotted for differentBPs in the inset.

is not without aws and approximations. Since the FSW process data is used to calculate

ow stress values the data set may be more realistic than Gleeble approach. Experiments

with large envelope may provide greater temperature window and greater condence to

calculated ow stresses.

Fig.4.29 shows the torque transients measured at the last 50mm of the welds made in

25.4mm thick AA 6061 using four different backing plates. The peak temperatures at two

locations of the probe in all the four cases were discussed previously. The average values

from the all the torque transients are shown in the inset with their respective standard devi-

ations as error bars. For the welds made using different diffusivity materials, whereas the

midplane peak T remains essentially identical, the root peak T decrease from low to high

diffusivity, the torque and hence power increase with increased backing plate diffusivity.

However this increase in the torque is associated with the increase in the forge force as

well (see Fig.4.18). Whether the increase in the torque is solely due to the forge force or

94

Property Workpiece Tool(Steel) Probe(MP159)Thermal Conductivity (W/m.C) 0.1265T+153.4 28 14.7

Heat Capacity(J/Kg.C) 0.8509T+825.7 490 421Density(Kg/m3) 2810 7750 8330

Table 4.1: Thermal properties and density of materials used in Simulation. Thermal prop-erties for the workpiece are expressed in terms of temperature and are for AA7075 adaptedfrom Guerdoux et al.[149] Other properties from [17, 150]

Temperature(C) Flow stress (MPa)25 450400 120425 100450 80475 60500 30532 0

Table 4.2: Flow stress-temperature relationship used in TPM model.[151]

the material used in the simulation are tabulated in Table.4.1 The temperature dependent

ow stress values used are shown in Table 4.2.

When the simulation is run, COMSOL calculates the steady state temperature eld by

solving the energy Eq. 4.2 in each element until satisfactory convergence is reached. The

3D geometry is meshed within COMSOL with tetrahedral element using boundary layer

mesh near the heat source edges such that an average mesh size of 0.5mm was attained

near the probe and shoulder interface where the gradients are known to be higher. Fig.4.36

shows some snapshot of meshed geometry with high mesh density near the heat source and

coarse mesh in the exterior.

As discussed in the section 2.5 on page 41 there are uncertainties in the thermal prop-

erties of the material used for simulation. Correct representation of the thermal properties

and boundary condition is important for accurate simulation of the process. It was thought

necessary to understand the qualitative effects of changing thermal properties and thermal

coefcients at the boundaries on the thermal eld. The understanding of effects of these

103

(a) Peak temperature effects

(b) Torque effects

Figure 4.37: Results of DOE analysis from Statgraphics. Main effect plot for peak temper-ature and torque shown on the left. Standardized Pareto chart on the right for each effect.

Having established that the contact conductance has the greatest impact on the peak

temperature achieved in FSW among different thermal coefcients, it is imperative to use

appropriate physically justied value. However, there is no agreement among different

researchers regarding its value. Although it is plausible that some variability in its value

exists depending upon the welding condition and thermal conditions, the differences prob-

ably should be within few orders of magnitude. However, a look at the literature shows that

researchers have use a wide range of contact conductance in their model ranging from 0.2,

200, 1000, 4000, 104,105 W/m2K (See Simulation review section for details). It is known

that contact conductance can depend on many local variables like applied pressures, tem-

perature, surface roughness and thermal properties of the materials involved which makes

it is difcult to characterize the values of contact conductance for each welding cases.

Nevertheless an attempt is made to assess upper and lower limit of contact conductance

106

using some experimental results and simulation. Experimental probe T obtained in 4.2mm

thick AA6056 by using different backing plates ( See Section 4.1) may provide some guid-

ance into reasonable value of effective contact conductance that may be applied during

modeling. For this purpose a series of TPM simulations were conducted by setting the

heat transfer coefcient at the WP/backing plate interface (h_bp) systematically at 7, 70,

700, 7000, 7104, 7105 and 7106 W=m2K. For each value of the coefcient the back-

ing plate conductivity were also varied being equal to that for aluminum, steel, AL6XN

and ceramic tile. The temperature attained at the midplane of the probe during from the

model are plotted in dashed lines against the backing plate conductivity in Fig.4.38 Also

plotted in red solid lines are experimental peak temperature data at different forge forces

made with various backing plates. These data sets were discussed previously in section

4.1. Considering the simulation data (dashed black lines) there is no appreciable differ-

ence in the probe T among various backing plate cases at the h_bp values of 7, 70, 700

W=m2K. Some realistic changes in the probe T among different BP is observed at h_bp

=7000W/m2K. Admittedly the trend of temperature evolution from is far from matching

the experimental trend (red solid lines). However it has to be recognized that the experi-

mental and simulation temperature are not directly comparable. Since the model assumed

sticking condition and the forge force is not included in the solution, the model temperature

probably is overestimated. Since the thermal contact between different materials like alu-

minum with aluminum will be different from that between say, ceramic tile and aluminum,

this adds more uncertainty to the data set. Nevertheless value of contact conductance thus

estimated can provide guidance in the order of magnitude sense. Looking at the simulation

data of h_bp =7103W/m2K and h_bp =7104W/m2K, it can be conjectured that for this

case realistic value of h_bp lies between 7000 and 70000 W/m2K

Contact conductance at the workpiece backing plate, h_bp =7103W/m2K established

using the above treatment was then used to setup the parametric simulation to assess the

capability of TPM model. Heat transfer coefcients and other boundary conditions used for

107

(a) Front view of the model showing different thermal boundary condi-tions applied

(b) Isometric view of model geometry showing additional thermalboundary condition.

Figure 4.39: Boundary conditions applied in the TPM model

experimental section 3.2 on page 53, it is difcult to represent the data vs some independent

variable. For this reason the results from simulation are plotted against the corresponding

experimental results. A diagonal line representing x=y is also shown for each graph. There

is a good agreement between the simulation and experimental probe temperature over the

range of weld parameter. Measured torque and torque calculated from the model are also in

fair agreement with each other as shown in Fig.4.42. Torque was obtained from the model

by boundary integration of quantity obtained by multiplying temperature dependent ow

stress with the radial distance of each element at the shoulder and pin interface. Although

the model appears to perform fairly well to predict peak temperature and torque it cant

109

Figure 4.40: Representative transverse temperature prole extractedfrom TPM model for 6.35 mm thick AA7050.

(a) In Air (hambient=10) (b) Under Water(hambient=104)

Figure 4.41: TPM probe temperature plotted against the experimentalpeak probe temperature. Diagonal line is x=y line.

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Figure 4.42: Comparison between simulated vs experimental tempera-ture trends across different parameter set.

predict the far eld temperature and similar accuracy.

In Fig.4.43 far eld temperatures from simulation are compared to the experimental

values. The comparison however is indirect. As discussed in the literature review, the lo-

cation of minimum hardness in precipitation hardening alloys occurs at HAZ region where

the peak temperature during the welding is around 350C. The distance of HAZ minimum

locations extracted from hardness traverses (discussed later) are plotted against the corre-

sponding RPM (shown as circles). With the increase in tool RPM there is a general trend

of increase in minimum hardness distance. Note that the last data point corresponding to

1000rpm shows shorter minimum hardness location compared to data point to the left. This

is probably because of the use of higher welding speed (see 3.2) compared to other cases.

The distance from the weld center where the temperature is ~350C were extracted from

corresponding TPM temperature eld and are shown by triangles. As per the aging kinetics

of precipitation hardening alloy these two data should be coincident. Although the trends

match, the data points dont match as much as the temperature and torque data discussed

previously; albeit the errors are reasonably small. This apparent mismatch might suggest

that although the considered model is capable of reasonably predicting the near eld tem-

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Figure 4.44: Steady state peak probe temperature vs.tool rpm for welds made with Densimet and NimonicTools.Welds made in 6.35 mm thick AA7050.

Figure 4.45: Measured steady state torque vs. the tool rota-tion rate for in air and under water welds made with Den-simet and Nimonic Tools. Welds made in 6.35 mm thickAA7050.

of 38107 m2/s. A series of welds was conducted in 6.35 mm AA7050 with tools made

of these two materials. The objective here is to understand the effects of varying thermal

gradient of the tool on the peak deformation zone temperature and other process responses.

Fig.4.44 shows the steady state peak temperature plotted against the tool rpm for In Air

(IA) and Under Water(UW) welds with two extreme cases of tool diffusivity: Densimet

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and Nimonic. Surprisingly, the peak probe T for a given parameter set for both the cases

are very close to each other even when the thermal diffusivity values of the two vary by a

factor of 7. With very low thermal conductivity, it was expected that Nimonic tool might

result in hotter nugget owing to lesser heat transfer from the tool compared to higher con-

ductivity Densimet. The measured steady state torque values for the two sets of tools are

also compared in Fig.4.45 The torque and hence power requirement for welds made with

Densimet tool is slightly greater than that for Nimonic tool. Because of its high thermal

diffusivity Densimet takes away the greatest amount of heat from the stir zone causing the

ow stress to increase which in turn will increase the torque and hence power. The differ-

ence however is not large. Realizing that the measured torque is a function of the shoulder

interface average temperature and forge force and recognizing that the commanded forge

force for each parameter set on both the tools were the same, the torque measured may

be used as a metric of the average interface temperature. Fig.4.45 shows that Nimonic

average shear interface temperature is slightly higher than that of the Densimet tool. It is

also interesting to note that the difference between Nimonic and Densimet tool torque is

higher in Underwater weld than IA ones. The heat extraction rate from the tool into water

compared to that in to air is more acute in high diffusivity Densimet than low diffusivity

Nimonic tool. Having stated that, the overall variation is not that large . From the discussed

result it may be concluded that the thermal conductivity of tool material does not make a

signicant difference in the deformation zone temperature in the given gage length under

considered situation. Underwater results suggest that this effect becomes larger when the

heat extraction rate is higher.

To address the issue of results in the peak probe temperature for different tools, TPM

models were run for the three different tools with widely varying thermal diffusivities.

Parameter set of 540rpm and 6.8mm/sec was used for the model. Fig.4.46 shows the tem-

perature prole along the Z direction (vertical line along axis of the tool) for all the three

tool material cases. The shank of the tool which is actually secured in the spindle by a col-

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Figure 4.46: Steady state temperature prole from the TPM model along the Z axis fordifferent tool materials. The tool picture on the right is shown as guide to the spatialtemperature distribution.

let, is made up of H13 steel and is simulated with the equivalent thermal property. Fig.4.47

shows the temperature maps across the probe and shoulder in transverse cross section. The

experimental temperature transients measured from TCs located at two different spots in

addition to that at the probe are shown in Fig.4.48 The gure shows temperature compar-

ison among Densimet and Nimonic tool from weld made at three cascaded parameter set:

400rpm -5mm/sec, 650 rpm-6.8mm/sec, and 800 rpm-6.8mm/sec. Both In Air and Under

water data are shown. As stated early the probe temperatures in both the cases are remark-

ably close in both the cases. Away from the probe at positions S1, S2 (not shown) and S3

the Densimet tool is hotter than Nimonic tool.

From both simulation and experimental results, it is clear that there is a steep thermal

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Figure 4.48: Experimental temperature measured at two spots (S2 and S3) in the tooladdition to the probe core are plotted against the weld time for both Densimet and Nimonictool welds. Three sets of parameters were used on the same plate in cascaded manner.400rpm -5 mm/sec, 650 rpm-6.8mm/sec, and 800 rpm-6.8mm/sec.

4.3 JOINT CHARACTERIZATION

Microhardness results and correlations

Previous sections sufciently demonstrate our ability to affect process response variables

such as required torque and weld zone temperature as well as, potentially, enabling pro-

duction of good quality welds with lower value of forge force by modication of thermal

boundaries. Of equal importance is the ability to use the thermal boundary condition to

inuence weld properties, mechanical and otherwise. In this section of the chapter weld

properties such as hardness, strength and grain size are correlated with weld parameters.

The results presented elucidate the effects of thermal boundary condition at the workpiece

on weld properties.

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Effects of forge force and backing plate

In following pages the effects of forge force and backing plate diffusivity on hardness

distributions of resulting welds are discussed using welds made in 4.2mm thick AA6056.

Shown in the Fig.4.49 are the transverse micro hardness proles for naturally aged samples

taken from welds made with a ceramic tile backing plate at three different forge forces for

the parameter set of 640rpm-6.8mm/sec. This set of graph illustrates changes in mechan-

Figure 4.49: Vickers hardness proles on transverse cross-section at midplane for naturallyaged samples. Weld performed at 6.8mm/sec using ceramic tile as backing plate. Forgeforce used and peak probe T reached for each cases are indicated. Welds made in 4.2mmthick AA6056-T451

ical property that occur in weld cross section with the change in peak temperature caused

by change in applied forge force. There is a signicant change in the shape as well as the

absolute value of the hardness distribution in the nugget with the change in applied forge

force. This change can be explained on the basis of the measured peak probe T during the

welding (indicated in the legend). As is evident from the graph nugget hardness increases

with the increase in the probe T. With a low measured peak T of ~398C obtained at a

low forge force of 5.7 kN, the weld nugget is in over-aged condition exhibiting a U shaped

hardness distribution. As the forge force is increased consequently increasing the peak

118

temperature, the nugget hardness starts increasing while the heat affected zones on both

the sides remain over-aged. Notice the nugget hardness asymmetry in weld made at inter-

mediate forge force. Several such cases have been observed in the current work and they

will be dealt in a separate section. With sufciently high forge force of 12.8kN, where the

measured peak T was recorded to be 490C, nugget hardness approaches base metal value

and asymmetry vanishes. This variety in nugget hardness at equivalent rpm and welding

speed can be explained by the different levels of peak temperature brought about by the

change in the forge force. Notice also that minimum hardness values for all the considered

samples located at HAZ for W shaped distributions and all over the nugget for U shaped are

in the narrow range of 70-74HV. This is reasonable and expected since thermal boundary

conditions and welding speeds for all the cases were identical such that time at elevated

temperature for precipitate coarsening has not changed with the change in forge force.

Fig 4.50 shows similar hardness results on welds made at 3.4, 6.8 and 10.2 mm/sec

respectively using high conductivity aluminum as backing plate material in contrast to

low conductivity ceramic tile that was discussed previously. Corresponding forge forces

used and peak Ts attained at the tool midplane are indicated in the legend. Consider the

(Fig.4.50a) where the hardness proles for welds made at 3.4mm/sec using three different

forge forces are shown. Despite a peak temperature increase from 392C to 431C there is

very small increase in the nugget hardness. All the three proles appear to be in overaged

condition. On the other extreme welding speed of 10.2 mm/sec (Fig.4.50c) for all the three

forge forces W shaped hardness is observed such that with higher applied forge force the

nugget hardness is higher and approaches the base metal hardness. The effect of achieved

peak temperature on the nugget hardness, all other things being equal can be appreciated

by comparing the hardness distribution among Aluminum (Fig 4.50(b) and ceramic tile

BP(Fig 4.49) at parameter set of 640rpm-6.8 mm/sec. With forge force of 8.5kN for in-

stance Aluminum BP sample has over-aged nugget with low hardness while with tile BP

the nugget has recovered some of its strength. The only difference between these two cases

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Figure 4.67: Percentage elongation and difference of hardness values between nugget andHAZ minimum plotted against the peak probe temperature. The dotted and solid lines aret by eye.

that deformation during transverse tensile testing tends to be concentrated in the HAZ. As

theDVHN becomes larger and larger, the strain concentration in the HAZ becomes more

prominent and the overall % elongation declines. Once theDVHN reaches a high enough

value, essentially no deformation occurs in the nugget during transverse loading: here the

critical value ofDVHN is between 35 and 40. The % elongation thus is not accurate a

reection of the ductility of the various weld regions since it is the weakest region that de-

termines how much stress is actually applied on the sample. The % elongation measured by

extensometer is the integral of the strain from the entire region where jaws of extensometer

were clamped.

140

(a) (b)

(c)

Figure 4.69: Average nugget grain sizes at different regions on welds made at 640 rpmand 6.8mm/sec using different forge forces on a) steel b) AL6XN c) ceramic tile backingplates.

backing plate. Measured grain size for welds made using 3 different backing plates at vari-

ous forge forces at 640 rpm and 6.8mm/sec including data shown in Fig.4.68 are plotted in

Fig.4.69. The grain size measurements for aluminum BP at this parameter set is not shown

since the grain boundaries were not revealed satisfactorily to allow sufcient condence

in grain size measurements. The effect of backing plate diffusivity in the grain size dis-

tribution is apparent from the graph. As the diffusivity of the backing plate decreases the

through thickness homogeneity in the grain size is enhanced. With Ceramic tile BP, the

grain size are about the same throughout the thickness. Grain microstructure from similar

situation as described above but for signicantly thicker (25.4mm thick AA6061) is shown

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Figure 4.74: Representative nugget microstructure and average grain size at crown, mid-plane and root for welds made at same rpm but different welding speeds as indicated.

sured peak probe temperature with corresponding welding speeds indicated by different

symbols in Fig.4.75. Notice that at the welding speed of 6.8 mm/s welds were performed

at three rotational speeds, hence three data points. The arrows indicate corresponding grain

sizes for in air and under water welds made using same sets of rpm and welding speeds.

The grain sizes of welds made below 400 rpm, with peak probe temperatures below 400C,

were not resolvable by optical means and are not reported here. For the rest of the weld

parameters, the underwater nugget grain size is consistently smaller than the corresponding

in air cases. This can be attributed to the lower peak temperature in the underwater welds.

For AA6056, the nugget grain size measured at the midplane from all the samples welded

at 6.8mm/sec and 10.2mm/sec are plotted against the measured peak probe T in Fig.4.76.

The grain size generally increases with the increase in the peak T at the nugget.

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Figure 4.75: Average nugget grain size measured at midplane plotted against thepeak probe temperature for in air and under water ambient cases. Different symbolsindicate different welding speeds. Welds made in 6.35mm thick AA7050-T7

Figure 4.76: Midplane nugget grain size plotted against the peak probe temperaturefor welds made in AA6056 using different backing plates and welding speeds asindicated.

148

made of AA6061. The welding conditions and parameters sets were otherwise identical

among the two cases: 480rpm-6.8mm/sec and forge force = 47 kN. The workpieces used

for this study were solution heat treated at 530C into a W/T4 state. This is different than

previously presented cases of welds in AA6061 where the initial temper of the work piece

was T6. As indicated in the legend of the gure the steady state peak probe T of the tool at

the midplane are similar among steel and composite BP. Near root temperature in compos-

ite BP is ~10C ( not a large difference) lower than that with steel BP most likely owing

to greater heat transfer with high diffusivity aluminum side bars. Note that with composite

backing plate the hardness traverses are much more close to each other at different thick-

ness levels suggesting better homogeneity in the hardness. In Fig.4.79 the average hardness

at nugget and HAZ minimum for different regions are shown. The nugget hardness at the

midplane and the root regions can be reasonably correlated to the measured peak T just like

previous results. Higher nugget hardness in steel BP compared to composite BP is probably

due to high level of solution heat treatment because of higher temperature exposure. Most

strikingly, the HAZ minimum hardness in the case of composite BP near the root region is

substantially greater than that with monolithic steel BP: 18VHN greater in advancing side

and 10 VHN greater in the retreating side.8 This increase can be decidedly attributed to

better cooling rate at the HAZ with aluminum side backing bar. TPM model described in

section 4.2 was used to qualitatively assess the temperature distribution in the weld made

with different congurations of the proposed backing plates. The simulation consisted of

a central strip of Ti-6-4 while the side bars were made out of copper. Five different con-

gurations were simulated with progressively narrower central strip of Ti-6-4 to assess the

changes in thermal prole as backing plate above the HAZ changes from low diffusivity Ti-

6-4 to high diffusivity copper. Fig.4.80 shows transverse contour plot of temperature with

progressively narrower central strip. With narrower Ti central strip the isotherms get more

8Note that even for steel BP the minimum hardness has signicantly increased ( from ~72HV comparedto welds made in T6 as opposed to T4 temper in this weld set; another reason could be that the work piecewas much wider in this case 610mm vs. 153 mm resulting in a greater thermal mass for heat sink.

151

inclined and hence elongated towards the horizontal direction. This probably indicates that

the use of composite backing plate changes the shape of the soft zone (~350C isotherm as

indicated by black dotted lines) thus may make the zone stronger in transverse tension. The

rate of cooling from 350C at the approximate minimum hardness region were extracted

from the models shown in Fig.4.80a ( Ti-6-4 BP over HAZ) and Fig.4.80e ( copper BP

over HAZ) and are plotted in Fig.4.80f. Although the simulation exaggerates the effect

because of the use of Ti-6-4 and copper backing plates in place of steel and aluminum in

the experiment, the signicant difference between cooling rates is in consort with the in-

creased HAZ minimum hardness observed in composite BP compared to monolithic steel

backing plate. The concept of composite backing plate has also proven useful to prevent

tool marks and indentation, and undesired lap weld at the backing plate which often occurs

when monolithic aluminum is used as backing plate.

153

CHAPTER 5

SUMMARY, DISCUSSIONS ANDCLOSURE

Various thermal management methods were applied during friction stir welding of three

avors of precipitation hardening aluminum alloys. Systematic body of knowledge per-

taining to the effects of thermal conditions at the backing plate, work piece surface and

tool on the process response and resulting joints were investigated. Following sections list

important conclusions drawn out of this work.

Conclusions from ambient condition modification works in 6.35 mm

thick AA7050-T7

In this study the effects of changing the thermal boundary condition at the surface of the

workpiece and control variables on weld properties and response variables have been ex-

amined. Several general trends have been observed and some useful correlations are found.

Specically:

1. All other things being equal increased surface convection resulting from welding

underwater compared to in-air welding results in:

a) Reduced nugget temperature.

b) Increased torque and power consumption.

c) Decreased average nugget grain size.

d) Increased cooling rates in the HAZ.

154

4. Use of novel composite backing plate consisting of low diffusivity central strip of

steel and side bars of aluminum backing plate has produced promising results. Trans-

verse micro-hardness results show that composite backing plate yielded a stronger

weld joint compared to weld made using monolithic steel backing plate. Improve-

ment is highest near the root of the weld.

Most of the correlations among various control parameters and boundary conditions with

process responses and joint properties presented in this work are consistent, physically

realistic and compelling. Nevertheless it is critical to acknowledge some limitations and

shortcomings. The effects of under water and sub-ambient conditions have been elucidated

for only a single base material and gage (6.35mm thick AA7050). AA7050 is known to be

highly quench sensitive. The degree of effectiveness of sub-ambient welding in other alloys

and gages will vary depending upon thermal and metallurgical variations. The welding

speed at which diminishing returns in HAZ minimum hardness improvement is reached

may also similarly vary. The effects of backing plate material and forge force in altering

thermal condition at the nugget and HAZ is well established from the results and analyses

of 4.2mm thick AA6056. However same has not been done effectively for thick plates. The

results from welding performed in 25.4mm thick AA6061 shows backing plate effects on

temperature, but because of the limited number of weld runs in AA6061, the effects have

not been explored fully. The effects of backing plate diffusivity and forge force have been

elucidated only for a single welding speed and rpm with a small range of peak temperature

unlike the work in AA6056 which covers a wider window. The relationship among toque,

tool forces and temperature are more difcult to deduce than in the case of relatively thin

AA7050 and AA6056. For thicker welds with longer probe, as the uncertainty in the shape

of the heat source and hence partitioning of heat source among the probe and shoulder

becomes more unclear, the relationship becomes murkier. The composite backing plate has

performed effectively for 25.4mm thick 6061. Nevertheless the use of composite backing

plate will be most fruitful in the case of difcult to weld alloy like thick AA7XXX series.

158