1. No category

Workpiece surface integrity when slot milling -TiAl
intermetallic alloy
Hood, Richard; Aspinwall, David; Soo, Sein; Mantle, Andrew L.; Novovic, Donka
DOI:
10.1016/j.cirp.2014.03.071
Citation for published version (Harvard):
Hood, R, Aspinwall, DK, Soo, SL, Mantle, AL & Novovic, D 2014, 'Workpiece surface integrity when slot milling TiAl intermetallic alloy' CIRP Annals - Manufacturing Technology, vol 63, no. 1, pp. 53-56.,
10.1016/j.cirp.2014.03.071
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Download date: 15. Oct. 2014
CIRP Annals - Manufacturing Technology 63 (2014) 53–56
Contents lists available at ScienceDirect
CIRP Annals - Manufacturing Technology
jou rnal homep age : ht t p: // ees .e lse vi er . com /ci r p/ def a ult . asp
Workpiece surface integrity when slot milling g-TiAl intermetallic alloy
Richard Hood a, David K. Aspinwall (1)a,*, Sein Leung Soo (2)a, Andrew L. Mantle (3)b,
Donka Novovic c
a
b
c
Machining Research Group, School of Mechanical Engineering, University of Birmingham, Edgbaston, Birmingham, UK
Manufacturing Technology, Rolls-Royce plc, Derby, UK
Turbines, Rolls-Royce plc, Derby, UK
A R T I C L E I N F O
A B S T R A C T
Keywords:
Titanium
Surface integrity
Slot milling
Slot milling is presented as a potential manufacturing route for aerospace component feature production
when machining g-TiAl intermetallic alloy Ti–45Al–2Mn–2Nb + 0.8 vol.% TiB2XD using 2 mm diameter
AlTiN coated WC ball nose end milling cutters. When operating with flood cutting fluid at v = 88 m/min,
f = 0.05 mm/tooth, d = 0.2 mm, maximum flank wear was 65 mm after 25 min. SEM micrographs of slot
surfaces show re-deposited/adhered and smeared workpiece material to a length of 50 mm. Brittle fracture
of the slot edges was restricted to <10 mm with sporadic top burr formation observed up to 20 mm. Cross
sectional micrographs of the slot sidewalls showed bending of the lamellae limited to within 5 mm.
ß 2014 CIRP.
1. Introduction
Manufacturing components from gamma titanium aluminide
(g-TiAl) intermetallic alloys with features that have ‘acceptable’
integrity is extremely difficult. This is mainly due to their low room
temperature ductility, which is quoted as being <2%. Despite this,
they are seen as potential replacement materials for nickel based
superalloy turbine blades, not least because alloy density is
typically 4 g/cm3,which is around half that of nickel alloys and
their operating ceiling extends up to 850 8C [1].
Milling is likely to be an important operation in the manufacture of blades and a significant proportion of published machinability research on these alloys over the last decade has focused on
this process [1–6]. Such studies have investigated the effects of
operating parameters and conditions on tool life and surface
integrity when milling g-TiAl alloys. The majority of cutters
assessed however have been >6 mm in diameter [2–6], with the
workpiece geometric shape simplified to rationalise testing,
although Beranoagirre et al. [2] does detail the machining of a
sample blade component but without disclosing the resulting
workpiece surface integrity. Additionally, while publications on
the micro milling of alpha/beta alloy Ti–6Al–4V can be found [7],
equivalent data for g-TiAl in respect of micro/mesoscale blade
feature production is sparse. The current research attempts to
address these aspects.
2. Experimental work
All milling tests were carried out on a Matsuura LX1, 3-axis high
speed machining centre rated at 200–60,000 rpm, with a maximum
* Corresponding author.
http://dx.doi.org/10.1016/j.cirp.2014.03.071
0007-8506/ß 2014 CIRP.
rapid feed rate of 90 m/min and a maximum spindle power of
5 kW. Cutting tools were held in MST Mizoguchi HSK shrink fit tool
holders. Where required, Houghton Vaughan Hocut 3380 cutting
fluid was supplied at a flow rate of 6 l/min. Trials in Phases 1 and 2
used
two
cast
workpiece
ingots
which
were
200 mm 100 mm 25 mm in size with composition, Ti–45Al–
2Mn–2Nb + 0.8 vol.% TiB2. These had a grain size of 200 mm and
bulk hardness of 330–350 HV and were surface ground on all sides.
Phase 3 work involved machining of representative cast component
features on similar but proprietary workpiece material and only
involved finishing operations.
Tool wear was measured using a Wild microscope and a
toolmakers table with digital micrometre heads giving a resolution
of 0.001 mm. Images of tool wear were taken using a Canon digital
SLR camera. Workpiece surface roughness was measured using a
Taylor Hobson Form Talysurf 120L and an Alicona IF G4 microscope
system. For workpiece surface integrity evaluation, selected
samples were hot mounted in edge retentive bakelite then
prepared using a grinder/polisher. Workpiece surfaces and
cross-sections were analysed using a Leica optical microscope
running Buehler Optimet software or a JEOL 6060 scanning
electron microscope (SEM) to assess the level of damage. Residual
stress measurements were undertaken using the blind hole drilling
technique. Difficulties can be encountered when using X-ray
diffraction techniques with g-TiAl due to the material’s tetragonal
phase and high levels of lattice strain, leading to broad, overlapping
reflections (line broadening) [4]. Each target site was prepared by
light abrasion (two passes of 400 grade SiC paper) followed by
thorough degreasing (acetone). Strain gauge rosettes type EA-06062RE-120 (Vishay Precision), were bonded to the workpiece
surface using Loctite 407 adhesive. Each rosette was drilled using a
PC-controlled 3-axis drilling machine. Fig. 1 shows the experimental setup.
54
R. Hood et al. / CIRP Annals - Manufacturing Technology 63 (2014) 53–56
Table 1
Operating parameters and conditions for Phase 1 tests.
Fig. 1. Schematic of residual stress measurement.
Depth increments were set at 32 mm (four times), 64 mm (four
times) and 128 mm (eight times), giving a completed hole depth of
1.41 mm for the determination of stresses up to 1024 mm below
the machined surface. Results from the individual target gauges
were obtained in the form of relaxed strains, which were
subsequently converted to stress values. Fig. 2 shows schematics/images of the coated cutting tools supplied by Seco Tools
(UK); WC XMG (WC 90% Co 10%, grain size 0.8 mm). Six different
tool formats were used associated with 5 basic types (denoted
Type 1 to Type 5), all of which were 2 mm in diameter and available
as stock items (except Type 5). As limited data was available for
small diameter end milling of g-TiAl alloys, it was decided to test a
variety of possible tools. The Type 1 tool was a 2 flute ball nose end
mill with a Mega 64 (AlTiN-Monolayer) coating. The cutting rake
was 08, helix angle was 178 and the cutting/flute length was 2 mm.
Two different tool subtypes; 1(a) and 1(b) were used with the
former having a maximum cutting depth of 4 mm while for the
latter it was 10 mm.
Fig. 2. Cutting tool types.
Tool Type 2 was similarly a 2 flute ball nose end mill with the
Mega coating (AlTiN) but having a cutting rake and helix angle of
108 and 308 respectively, while the cutting length and maximum
cutting depth was 9 mm. The Type 3 cutter was a 2 flute high feed
end mill with a Mega coating having a corresponding cutting
length of 0.15 mm and maximum cutting depth of 6 mm. Tool Type
4 was a 4 flute end mill with a Mega coating. The cutting rake was
28, helix angle was 208 while the cutting length and maximum
cutting depth were 2.5 and 5 mm respectively. The Type 5 tool was
a bespoke design produced specifically for the current work, which
was based on the Type 1 ball nose end mill with a cutting length
and maximum cutting depth of 9 mm. This extended cutting/flute
length was designed to reduce material smearing.
Experimental work involved 3 phases, which employed down
milling unless otherwise stated. In Phase 1, research was aimed at
determining tool life and workpiece integrity performance when
machining the sidewall of a slot under various finishing conditions.
Table 1 details the variable test parameters utilised with the radial
depth of cut/stepover fixed at 0.2 mm. The levels chosen were
based on tool manufacturer recommendations and previous
research using larger diameter tools [4]. Tests 1 and 2 were
performed using Type 1(a) tools to compare dry and flood coolant
operation. Tests 3 and 4 were carried out to investigate the effect of
tool geometry (Type 1(b) vs. 2) on tool life, with operating
parameters selected based on results from Tests 1 and 2. Tests 5
and 6 investigated the effect of cutting speed to determine the
maximum possible value before tool failure/heavy wear occurred
using Type 3 cutters.
Test
Tool
type
Cutting
speed
(m/min)
1
2
3
4
5
6
7
8
9
10
1(a)
1(a)
1(b)
2
3
3
4
4
4
4
55
55
88
88
55–160
160
88
88
88
88
Feed
rate
(mm/tooth)
Axial
depth of
cut (mm)
Cutting
environment
0.01
0.01
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.05
0.1
0.1
0.2
0.2
0.2
0.2
0.1–1.0
0.05
0.2
0.05
Dry
Wet
Dry
Dry
Dry
Dry
Dry
Wet
Wet
Wet
Test 7 involved tool Type 4 to investigate the maximum depth
of cut before tool breakage. Here, varying axial depths of cut from
0.1 up to 1.0 mm were employed with the maximum machining
depth of 1.0 mm. Test 8 provided an assessment of cycle time and
compared workpiece surface integrity when using three different
milling strategies namely single direction rasters; one providing
down milling, the other up milling and conventional raster
operation giving alternate up and down milling scenarios.
Operating parameters were similar to those in Test 7 but with
flood coolant and a 0.05 mm axial depth of cut. Residual stress
measurements were undertaken on surfaces machined using Test
8 parameters with a raster milling strategy along with a 0.2 (Test 9)
and 0.05 mm (Test 10) depth of cut. Financial constraints limited
additional measurements, the two tests chosen giving ‘preferred’
parameters for workpiece integrity in Phase 1 work.
Phase 2 work involved the milling of an ‘end slot’ from solid
(roughing). The slot was 2.5 mm wide, approximately 12 mm long
and 9 mm deep. Several tests (not all reported here) were
performed using cutting parameters based on Phase 1 results to
determine their effect on tool life and workpiece integrity. Fig. 3
shows a schematic with cutter paths developed using Delcam
PowerMill software (Fig. 3(a)) and a section of the machined
workpiece (Fig. 3(b)). A profile cutter path was used. Tests 1 and 2
compared the number of end slots machined against flank wear for
Types 1(b) and 3 tools using a cutting speed of 88 m/min and a
depth of cut of 0.05 mm with flood coolant. The feed rate for the
Type 1(b) tool was 0.05 mm/tooth while for the high feed end mill
(Type 3) it was and 0.1 mm/tooth.
Fig. 3. Phase 2 workpiece (a) CAM model, (b) workpiece.
As outlined previously, Phase 3 research involved machining of
representative features on ten cast workpieces (finishing only).
Commercial restrictions preclude detailing the exact geometry
however workpiece surface integrity data is presented. Machining
parameters were based on the preferred values from Phases 1 and
2, including a cutting speed of 88 m/min, a stepover of 0.2 mm and
flood coolant. Variable parameters were feed rates of 0.025 and
0.05 mm/tooth along with axial depths of cut of 0.05, 0.1 and
0.2 mm. A new Type 5 tool was used for each workpiece.
3. Results and discussion
Phase 1 experimental results for Tests 1, 2, 3, 4 and 6 are shown
in Fig. 4. Tool life for Tests 1 and 2 was generally good with a
R. Hood et al. / CIRP Annals - Manufacturing Technology 63 (2014) 53–56
Fig. 4. Tool life curve for Phase 1 sidewall milling.
machining time in excess of 100 min using Test 1 operating
parameters. Plotted data are the averages for the maximum flank
wear values on individual teeth. Results for Tests 1 and 2 were
comparable until a machining time of 60 min after which Test 2
(involving use of flood coolant) showed a higher wear. Given the
limited literature available on slot milling of titanium alloys with
small diameter tools, the parameter levels selected for these tests
were conservative. Test 3 showed a higher rate of tool wear, yet
despite the 60% increase in cutting speed (55–88 m/min), the
average maximum flank wear level was only 60 mm after 25 min.
In contrast, the cutter in Test 4 fractured on first contact with the
workpiece, which is not an uncommon occurrence when micro
milling titanium alloys [7]. The rake angle for the Type 2 cutter
used in this test was +108 compared with 08 for Test 1–3 tools.
Priarone et al. [3] describes a similar result with a 50% increase in
tool life when comparing tooling with a radial rake angle of +128
and +48. Test 5 showed a maximum cutting speed of 160 m/min
was achievable and therefore Test 6 was performed as a life trial
using this value. Tool life was however extremely short, with a
machining time of only 30 s for an average maximum flank wear
level of 200 mm. Wear scar photographs from Test 1 and Test 6 at
cessation are also shown in Fig. 4.
Tool breakage did not occur even when using a depth of cut of
1 mm with the Type 4 tool in Test 7 and five passes were achieved
before the test was stopped. Test 8 showed only marginal
differences in surface/subsurface damage after machining in either
a single raster up or down milling mode, or when raster milling
with alternate up/down operation. No cracking was observed and
bending of lamellae was restricted to <4 mm. Consequently, the
reduced cycle time associated with conventional raster operation
was deemed preferable.
Residual stress depth profiles for Tests 9 and 10 (new tools) are
detailed in Fig. 5. Both trials showed similar trends with a
maximum compressive stress of between 400 and 527 MPa at the
surface, which became neutral at a depth of 450 mm. The lower
depth of cut for Test 10 produced a reduction (25%) in residual
stress for each measurement. In both tests, the residual stress
parallel to the feed was 10% higher (more compressive) than that
measured perpendicular to the feed direction. Measurement
Fig. 5. Residual stress depth profile measurements.
55
uncertainty was 85 MPa at a depth of 16 mm, 19 MPa at a depth
of 512 mm and 23 MPa at a depth of 1024 mm. These errors consist
of a random strain uncertainty of 3 me and a linear error of 7% of
the total stress level [8].
Published residual stress data for ball end milling g-TiAl using
6 mm diameter cutters, details compressive residual stresses
ranging in magnitude from 600 to 1200 MPa [4]. The lower levels
with the current work reflect the less arduous operating
parameters used, yet both sets of results concur that the stress
became neutral at a distance of 0.3–0.5 mm below the workpiece
surface. In general, cutting temperatures when ball nose end
milling g-TiAl alloys are low, of the order of 200–300 8C for a
cutting speed 70–120 m/min, feed rate 0.12 mm/tooth and depth
of cut (axial and radial) 0.2–0.5 mm [5]. Operating regimes subject
to low cutting temperatures coupled with a high strain hardening
rate, typically produce compressive stresses which are likely to
produce longer fatigue life as the peak tensile load at the surface
would be reduced and surface cracks would tend to close [4].
Workpiece surface roughness (Ra) measurements performed on
Test 9 and 10 samples showed a range in values from 0.8 to 1.2mm.
Phase 2 work machining slots from solid as detailed in Fig. 3,
produced mixed results. For the high feed tool (Type 3), 25 slots
were machined before the average maximum flank wear level of
200 mm was reached, while the ball nose end mill (Type 1(b))
produced 10 slots before tool edge fracture occurred and the test
halted. Fig. 6 shows the wear progression with a wear scar
photograph for the ball nose end mill trial at test cessation.
Fig. 6. Tool life curve for Phase 2 solid slot milling.
Images of the machined workpiece surface from Phase 2 are
shown in Fig. 7. The surfaces were covered extensively in smeared/
adhered workpiece material with a length of up to 50 mm, see
Fig. 7(a and b) for the slot mid point and closed end respectively.
Possible cracks with a length of up to 100 mm, were also found as
shown in Fig. 7(c). It was often difficult to ascertain if these were
cracks in the bulk workpiece or adhered material. Additionally
edge fracture was observed, see Fig. 7(d), despite the low level of
operating parameters, edge profiling would be required in practice
Fig. 7. SEM images of the Phase 2 machined workpiece surface.
56
R. Hood et al. / CIRP Annals - Manufacturing Technology 63 (2014) 53–56
to minimise this. Results showed considerably worse surface
integrity than found in Phase 1.
Scanning electron microscope images of the cast workpiece
surfaces machined as part of Phase 3 testing showed surface
defects on nearly every cast feature. These were less severe than
those detailed for Phase 2 work and were consistent with Phase 1
finishing investigations. The damage consisted of tearing of the
surface and fracture of the milled/cast surface interfaces as shown
in Fig. 8. In general, the size and incidence of the tearing increased
as the level of operating parameters increased. As with the residual
stress results, the level of workpiece surface damage was
somewhat less than that reported in previous research with larger
cutters and higher material removal rates [6], particularly with
respect to fracture/pullout. Having said this, a large amount of
smeared/adhered workpiece material was observed in the SEM
images, which may have obscured any surface damage, see Fig. 8(a
and d).
similar cracks were seen. It is likely therefore that they were a byproduct associated with the casting process. Not surprisingly, tests
performed under more arduous conditions which included a depth
of cut of 0.1 mm and feed rate of 0.05 mm/tooth, showed greater
bending of the lamellae and an ‘arc’ shaped crack (6 mm deep and
80 mm wide) was observed as shown in Fig. 9(c).
Cross-sectional analysis of the cast/milled surface interfaces
showed mixed results. For some cast features there was limited
damage (<10 mm), whereas for others more damage was
observed. Sporadic top burr formation as shown in Fig. 9(d)
occurred on several cross-sections with dimensions of up to
20 mm. Burr formation when machining brittle materials like gTiAl is relatively rare but presents an additional problem requiring
an edge profiling solution. However, the sporadic nature of the
formation is likely to preclude the use of a prediction system to
modify the design to avoid burrs [9].
4. Conclusions
Fig. 8. SEM images of the machined surface of the cast workpieces.
The majority of workpiece cross-sections produced following
machining of the cast features, also showed some degree of
damage; see Fig. 9. Where visible this consisted of bending of the
lamellae to an average depth of <5 mm and a maximum depth of
15 mm as illustrated in Fig. 9(b). Cracks (100 mm long and 4 mm
wide) between lamellae were observed in two locations within
0.5 mm of each other on workpiece 3 (from the batch of 10). This
was the only time that cracks of this nature were observed, indeed,
in two further replications using identical operating levels no
Fig. 9. Cross-sectional micrographs for Phase 3 investigation.
Workpiece surface integrity characteristics including residual
stress data, have been established for the slot milling of
intermetallic alloy Ti–45Al–2Mn–2Nb + 0.8 vol.% TiB2 when using
small diameter end milling tools. Such data is a prerequisite in
identifying appropriate strategies for the manufacture of aeroengine components. Residual stress measurement using the blind
hole drilling technique showed compressive residual stresses of
527 MPa which became neutral at a depth of 450 mm. Slot
surfaces showed re-deposited/adhered and smeared workpiece
material to a length of 50 mm. Brittle fracture of the slot edges
was restricted to <10 mm with sporadic top burr formation
observed up to 20 mm. Slot sidewall micrographs detailed
bending of the lamellae limited to within 5 mm when using
finishing conditions. Slots machined from solid showed the worst
surface integrity of all the features machined.
Acknowledgments
We would like to thank Rolls-Royce plc and Seco Tools (UK) Ltd.
for both financial and technical support.
References
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