Unbalanced tooling can introduce considerable detrimental effects on not only the machine tool – this high centrifugal force causing internal bearing stresses leading to premature spindl
Trang 1with the rotational axis, but intersects it at the
cen-tre of gravity of the ‘assembly’s body’ Under such
conditions the force vectors equalise, but are 180°
apart
3 ‘Dynamic unbalance’ – dual-plane Such a
condi-tion of the toolholder assembly arises when the axis
does not coincide with the rotational axis and is not
either parallel to, nor intersecting this axis (i.e see
Fig 232)
For any rotating tooling assembly, estimating the
cut-ter unbalance is possible using the following variables:
M = cutter/holder mass,
S = mass centre,
e = displacement of mass centre,
r = distance from centre of tooling, to the centre of
gravity of mass (m),
ω = angular velocity,
m = mass unbalance,
U = cutter unbalance,
9549 = a constant
Determining the relative unbalance (U) of a
rotat-ing toolrotat-ing assembly, can be found by the followrotat-ing
expression(s):
U = M × e or, alternatively: U = m × r (i)
It is usual to express unbalance in terms of the product
of the mass times distance, typically using the units:
‘g-mm’
Finding the magnitude of centrifugal force produced
by the rotating tooling assembly with a given
unbal-ance, can be established as follows:
Where: ‘ω’ is the angular velocity in units of radians
sec–1
The formula to find ‘ω’ is expressed by:
Therefore, by combining formulae: (i) and (iii), in (ii),
we can obtain the magnitude of centrifugal force ‘F’ , as
follows:
As established in equation (iv), the centrifugal force
caused by tooling unbalance will increase by the
‘square of the speed’ , in a similar manner to the spin-dle nose taper swelling (i.e growth) previously men-tioned Nonetheless, assuming that this specific tool-holder initially has a low unbalance, this will become a problem if the rotational speeds are increased beyond 10,000 rev min–1 For example, with most toolholders
exhibiting single-plane unbalance, research
experi-mentation has shown that the initial unbalance of a
typical tooling assembly will be of the order: 250
g-mm When such tooling is rotated at 15,000 rev min–1, this 250 g-mm of out-of-balance develops a continu-ous radial force of 642.6 N
Unbalanced tooling can introduce considerable detrimental effects on not only the machine tool – this high centrifugal force causing internal bearing stresses leading to premature spindle failure, but affects cut-ter life and degrades workpiece surface texture Much
of the principal tooling unbalance problems can be traced-back to several sources, such as:
• Toolholders of the V-flange type, which might have different depth of drive/slots, these toolholder fea-tures being part of the inherent design,
• Toolholders for some end mills and slot-drills, hav-ing set screws for lockhav-ing the cutter securely in place, so due to necessary clearance and the radial application of the set screw, this creates minute cut-ter eccentricity – causing unbalance,
• Out-of-balance caused by an unground V-flange base,
• Collet and its collet nut tend to be recurring sources
of unbalance in HSM tool holders
NB Most of these tool holding-related issues can
be eliminated by simply modifying the tooling de-sign
As can be seen from Fig 232, the marginally eccen-tric adjustable balance rings can be rotated to adjust the degree of single-plane balance, with several of the tooling manufacturers offering differing adjustment methods for HSM toolholders
Finally, consideration needs to be given to the level
of balance-quality required and in HSM applications for example, a milling cutter is expected to withstand
‘Single-plane unbalance’ , relates to the type of unbalance that
occurs in either one of two planes Namely, the tooling
assem-bly’s single-plane unbalance will be in either its axial, or radial directional plane.
Trang 2both high rotational speeds and associated cutting
forces, thus here it can be considered as a ‘rigid
ro-tating body’ This assumption allows one to use the
ANSI S2.19-1989 Standard, for achieving balance
– see Fig 233, which defines the permissible residual
unbalance of a rotating body relative to its maximum
speed This Standard and its equivalents (e.g ISO:
1940:1; ISO: 1290 G), assigns different balance-quality
grades termed: ‘G-numbers’ , related to the grouping of
rotating bodies (i.e not shown), these groupings
be-ing based upon the experienced gained with a variety
of: sizes; speeds; and types Thus, the balance-quality
grade ‘G’ , equals the specific unbalance ‘e’ times the
rotational speed ‘ω’ , as follows:
Balance-quality G = e × ω (mm sec–1)
Furthermore, the equation was described earlier, thus:
∴ solving for ‘U’ , we obtain:
From the Standard, the balance-quality for machine
tool drives is given as: G2.5, although in many
in-stances the value utilised should ideally approach that
of G1.0 – this being the specification for grinding
ma-chine tool drives, as today in HSM applications they
are compatible However, if for the purposes of
clari-fication of the unbalance tooling condition the value
of G2.5 is utilised, then the following worked example
illustrates the balance-quality necessary using a
tool-holder weighing 3 kg, rotating at 25,000 rev min–1:
U (higher)= � � ., (g-mm)
∴ U (higher) = 2.85 g-mm.
As alluded to previously, this unbalance condition
is the ‘worst case’ and the tooling should ideally
ap-proach G1.0, this balance-quality value, gives:
U (lower)= � � ., (g-mm)
∴ U (lower) = 1.14 g-mm.
This then follows that the balance is between 1.14 and
2.85 g-mm, which is toward the ‘upper-end’ for the
maximum residual specific unbalance for the G2.5, while approaching this level for the G1.0 (i.e shown by the graph in Fig 233)
Even when the tooling assembly has been dynami-cally balanced in both planes (i.e see Fig 234a – more
to be said on this topic shortly in Section 9.5.2), prob-lems still exist, particularly in the fit of the spindle ta-per connection (Fig 232) This is a result of the tata-per rate accuracy requirements between both the shank and taper socket In fact, the situation is quite a con-fusing one, due to the relative cone ‘Angle Tolerance’ grades: AT-1 to AT-6, that are employed using the con-ventional fitment of: 7:24 taper Not only do different countries often have their own connection Standards, but previously, even individual machine tool manu-facturers within each country had adopted differing Standards! Today, many machine tool companies tend
to utilise taper spindle connections that are compat-ible to an appropriate Standard and complement those
of the tooling manufacturers
9.5.2 HSM – Dynamic Balancing
Machine Application
It has been discussed in the previous sections that cut-ting tool assemblies when combined with an HSM strategy, can be a large contributor to dynamic unbal-ance For instance, in the production and manufacture
of say, the geometry of a face-mill, the tooling stock material is: externally/internally turned on one side; unclamped; flipped-over and rechecked; then turned
on the other side; then located onto a milling machine tool for operations on the individual insert pockets that must be milled; and indexed – as appropriate for the number of cutting edges; this necessary clamp-ing/reclamping workpiece (i.e face-mill) procedure,
will create a tool that is marginally-unbalanced With
HSM, the otherwise unnoticeable unbalance at con-ventional rotational speeds, becomes intolerable in these high-speed ranges Often, the most economical technique for achieving balanced tooling for tooling
‘Insert pockets’ , are sometimes ‘differentially-pitched’ which
means they have unequal spacing of teeth around the cutter’s periphery This pitching technique for cutting insert pockets,
is quite effective as a means of reducing machining vibrational effects often encountered with coarse-pitched face-mills.
Trang 3Figure 233 A graph to determine high-speed cutter unbalance ‘U’ (ANSI S2.19–1989)
.
Trang 4designated for the HSM ranges, is by dynamically
bal-ancing the tools in an appropriate machine (Fig 234a)
Thus, during tool balancing, the cutter is clamped in
a fixture that rotates in very rigidly supported
bear-ings (Fig 234a) Any unbalance in the rotating
cut-ter is directly measured as centrifugal force, which is
transmitted along with its actual rotational position
to a specially-configured computer The computer
calculates the amount and location of the material to
be removed from the tool’s body, in order to properly
adjust the mass distribution This unwanted tool
ma-terial can then be removed, by either drilling holes,
or milling flats (i.e see Fig 234b – illustrating that a
small amount of material has been removed – milled –
from its flange and is termed ‘Hard-balancing’)
Usu-ally, flats are preferred for removing larger amounts of
tool material stock, because at high rotational speeds a
tool with a hole can generate a unacceptable ‘whistling
noise’: see Fig 235 to gain an appreciation of the effect
of these high peripheral speeds, here, a large face-mill
is shown in an HSM automotive application Since
the ‘balancing operation’ can compromise the tool’s
strength and performance, it is important to establish
where any superfluous tool material can be safely
re-moved from on the cutter’s body
Balance is a ‘zero-quantity’ , so it is customary to
measure balance in the absence – within acceptable
limits – of unbalance Thus, the ‘balance tolerance’ is
the maximum residual unbalance (g-mm) allowed for
a particular tool’s weight and rotational speed For
ex-ample, the ANSI Standard quality-grades for balance
tolerance range from G0.4 to G6.30, with the lower the
‘G-number’ the closer the balance tolerance It should
be emphasised, that only when a balanced tool and its
balanced toolholder are balanced together as a
com-plete unit, are they truly dynamically-balanced
‘Dynamic-balancing’ postscript: the cutting inserts, screws
wedges that are retained in the cutter’s pockets must be
se-curely locked into position If these small items become
de-tached when HSM, this may cause disastrous consequences
for any operator in the vicinity Therefore, machine tool
guard-ing of more than adequate protection is vital here, to minimise
potential safety hazards to the personnel when in use.
9.6 HSM – Research
Applications
Introduction
Possibly the foremost reason for conducting applied re-search programmes at various universities and similar research-based organisations, is to ‘push the boundar-ies’ of our theoretical and practical understanding of current, or novel machining concepts Rather than at-tempt to give a perhaps less-than-informed account of what is transpiring at other ‘learned organisations’ It was decided just to deal with some recent work that was undertaken by the author, in association with in-dustrial and academic colleagues, both in the UK and abroad As these particular research projects were ei-ther undertaken, or instigated by the author, often in close union with others, mainly concerned with indus-trial-sponsored doctoral programmes, it was felt that here at least, some continuity concerning HSM-related research themes could be achieved
9.6.1 Ultra-High Speed: Face-Milling
Design and Development
Introduction
In order to achieve a cutting speed of say, 1,000 m min–1, a φ10 mm milling cutter has to rotate at ap-proximately 32,000 rev min–1 In fact, this is quite fea-sible for smaller diameter cutters, as there are quite a few of today’s machine tools having HSM spindles that can exploit these speeds However, it is worth a sight digression here, prior to continuing with the theme concerned with special-purpose UHSM by
face-mill-ing, to ask the pertinent question: ‘What do we mean
by HSM?’ Is it:
• High rotational speed machining?
• High cutting speed machining?
• High feed machining?
• High speed and feed machining?
• High productivity machining?
Even these five potential contenders for what amounts
to HSM, are by no means an exhaustive list, one could also add more obscure factors such as: the power de-mand on the spindle; tool assembly balance speeds;
Trang 5Figure 234 Dynamic dual-plane (i.e radial and axial) cutter balancing [Courtesy of Ingersoll]
.
Trang 6Figure 235 High-speed large diameter face (finish) milling on a grey cast iron engine block [Courtesy of Sumitomo
Electric Hardmetal Ltd.]
.
Trang 7the taper size in relation to its rotational speed;
thin-walled machining capability; etc.; the list will grow,
depending upon what we consider to constitute as
go-ing either very fast, or what speed allows us to machine
certain types of part features! Although even here, the
major benefits of an ultra-high speed machining
strat-egy are somewhat lost if a ‘working definition’ is not
clearly stated In this current discussion on the
sub-ject, one measure of HSM could be if the peripheral
speed of the cutter, or workpiece is >1,500 m min–1
Hence, this could represent a base-line for the
tran-sition from conventional to HSM, so for comparison,
as in the case for previously mentioned small
diam-eter milling cutter of φ10 mm, it would need to rotate
>47,750 rev min–1, or conversely, a larger face-mill of
say, φ300 mm, would have to rotate at approximately
1,600 rev min–1 – in order to sustain a peripheral cutter
speed of at least 1,500 m min–1 This latter rotational
speed although considerably slower, is a much greater
problem that that presented for the former small
diam-eter cutter The reasons for this are three-fold: firstly,
has the machine tool got enough spindle power to
achieve the necessary stock removal rates required, or
will it be likely to stall? Secondly, is the spindle taper
fitment robust enough to cope with the torque effects
and bending moments imparted during machining?
Thirdly, will the cutting inserts still be retained at the
high centrifugal forces generated in association with
and exacerbated by the imparted cutting forces? These
and other lesser important questions and decisions
need to be addressed, if the large diameter face-mill
is to successfully mill across a wide workpiece surface
with any degree of efficiency This former point of the
manner in how workpiece surface stock is removed is
important, for example, two markedly differing
ma-chining strategies could be adopted, such as:
I Shallow depths of cut combined with rapid
tra-verse rates and small step-overs, utilising smaller
diameter cutters at high peripheral speeds,
II Deep and wide cuts with a large diameter
face-mill with slower traverse rates, having much
lower rotational speeds
NB Both machining strategies will remove
simi-lar amounts of part stock!
UHSM: Face-Milling Cutter Design
When designing large diameter face-milling cutter
assemblies for production applications in the UHSM
range, a number of critical features need consideration, such as: cutter-body material; its weight and rigidity; taper fitment; dual-plane balancing; as well as its aero-dynamic behaviour – at fast peripheral velocities
If one attempts to design and develop a large face-milling cutter with an insert cutting circumference designed to rotate at 3,000 m min–1, which at first glance, may not seem that fast However, if we equate this cutting speed to that of the same φ10 mm mill-ing cutter previously mentioned, then this smaller tool would have to rotate at ≈95,500 rev min–1, but for the larger diameter cutter, it would also require
to be dual-plane balanced Without dynamic bal-ancing, the large cutter may be prone to a disastrous series of vibrational problems, which may ultimately lead to premature cutter failure – with all its attendant safety hazards The cutter design in Fig 236 for this applied research programme, was dual-plane balanced
to Standard defined in ISO:1940/1, being to a specific
‘G-number’ This Standard was initially conceived for the rotational balancing of impellers and similar high-speed equipment, across a large speed range The large face-mills had been dual-plane balanced to G2.5
@ 10,000 rev min–1 As mentioned in Section 9.5.1, this
‘G-number’ refers to the maximum tolerable imbal-ance for the complete tooling assembly, being based upon the previously described formula (i.e given be-low again for clarification), resulting in the folbe-lowing calculations:
Unbalance: U = � M � G N (g-mm)
[or] Force: F = U � N�(N)
Where:
U = allowable unbalance (g-mm);
9549 = a constant;
M = mass, or weight of the total cutter assembly
(kg);
G = pre-selected balance tolerance number from
ISO: 1940/1;
‘Dual-plane balanced cutters’ , are those cutters that have
been dynamically balanced in two planes: having both radial and axial balance.
Trang 8N = maximum rotational speed of tooling assembly
(rev min–1);
F = force (N)
When the above equation for cutter unbalance is
utilised, for these tooling assemblies, utilising the
fol-lowing values: G = 2.5; M = 4.294; N = 6,000; which
then gave an allowable imbalance U = 17.9 g-mm This
level of imbalance means that the cutter’s mass
can-not rotationally shift by more than 17.9 g-mm, if it is
to maintain dual-plane balance at a peripheral speed
of at least 3,000 m min–1 In fact, two identical cutter
assemblies were designed and manufactured,
hav-ing a BT40 taper fitment – for the vertical machinhav-ing
centre (i.e Cincinnati Milacron Sabre 500) In order
to maintain both structural rigidity and integrity,
the complete cutter bodies and their associated
ta-pers were each produced from single stock of EN24T
steel After precision turning and milling the complete
bodies and insert pockets, these cutters were
nitride-hardened to HRC 52, prior to a ‘very light-grinding’
process and then balancing The four cutting insert
pockets were equally-spaced (pitched) and the
but-ton-style cemented carbide inserts were: φ12 mm by
4mm thick, single-sided and TiN-coated (Stellram:
RPET 1204 DFZ) The insert pocket geometry had a
11° toe angle with a neutral geometry Button inserts
were selected as they give the strongest shape cutting
geometry available (see Fig 23), producing an infinite
approach angle to the workpiece (see Fig 83b), thus
minimising impact load at entry to the cut while
of-fering multiple cutting edges – when subsequently
turned in their seatings As one might expect, insert
security is vitally important, due to the great
centrifu-gal effects and applied cutting forces Due to the
pre-vious nitride-hardening process, hardened insert seats
were unnecessary, once the retaining screws had been
‘torqued-up’ locking and then sealing them – for
se- ‘Nitride-hardening’ , produces a very hard surface with a
softer and tougher matrix The UHSM cutters were held in a
pressure-tight furnace and heated to between 500-550°C for
some hours in an ammonia gas*, allowing the nitrogen atoms
to diffuse into the surface and to form fine stable nitride
pre-cipitates with aluminium constituents, allowing the
nitrided-steel surface to be precipitation- hardened No subsequent
heat-treatment is necessary.
*Approximately 30% of the ammonia disassociates ( NH ←→
3H+N) and part of the nascent nitrogen is absorbed by the
surface layers of the steel (Source: Cotrell et al., 1979)
curity The actual seatings for the inserts had consid-erable body-support around their periphery, just hav-ing a workhav-ing-clearance at the insert’s cutthav-ing region These UHSM face-mills were extremely compact, with the minimum stand-off height from the cutter’s gauge line (i.e see Fig 236), which reduced the effects of the previously mentioned ‘rigidity rule’
Due to the relatively large diameter and weight of these face-mills and the fact that the machining centre had limited spindle power, these cutters could, if used appropriately, exploit the ‘mass’ , or ‘flywheel-effect’ of their weight in conjunction with rotational speed to
‘store inertia’ So, when the spindle power is restricted, cutters with high mass must be taken up to their de-sired rotational speed in a progressive manner, other-wise they are likely to ‘trip’ a ‘spindle over-load’ in the CNC controller This steady and progressive increase in the cutter’s rotational speed occurred at 500 rev min–1
increments – dwelling for several seconds to minimise inertial power overload, between increases to the de-sired peripheral speed Due to the machine tool hav-ing a maximum spindle speed of 6,000 m min–1, this equated to a peripheral cutting speed of 3,000 m min–1, with the face-mill having 0.5 m cutting circumference Once the cutter has reached its top speed, it can then
be rapidly progressed (i.e fed) across the workpiece at
a rate of 20 m min–1 In this case the workpiece ma-terials were a range of stainless steel alloy testpieces (Fig 236) After rapidly face-milling these ‘stainless testpieces’ , the cutter’s rotation was decremented in
500 rev min–1 intervals until stationary The cutter once stationary, could have its edge wear (inserts) assused and workpiece milled surface texture and surface in-tegrity could be inspected and investigated
One factor to bear in mind concerning UHSM with large face-milling cutters being utilised for their ‘in-ertial effect’ , is to design them without driving dogs
If these ‘dogs’ were fitted, not only can they introduce out-of-balance effects, but tend to significantly disrupt the air-flow and introduce alarming and high noise factors This aspect of cutter design is important, if the cutter cannot ‘cleave through the air’ with aerody-namic efficiency, turbulent air flow will result and op-erational noise becomes excessive One problem that these particular cutters did not suffer from (i.e despite the conventional taper-cone angle) – unlike many of their higher rotational speed counterparts, was ‘spin-dle nose swelling’ , which can cause a lack of regis-ter if the taper fitment connection is not of either the double-, or triple-contact face-and-cone types One unexpected aspect of employing such large face-mills
Trang 9Figure 236 A specially-designed dual-plane (radial and axial) face mill, for ultra-high-speed milling [Source: Smith,
Wyatt & Hope, 1998]
.
Trang 10in UHSM, was found to be due to the very high
peri-pheral speed A ‘suction-effect’ resulted from the
un-derside clearance of the cutter’s body, created a
‘low-pressure region’ This ‘virtual vacuum’ here, meant that
conventional-pressure flood coolant application was
not possible, as it simply vaporised to a mist!
UHSM: Cutting Trials
The neutral geometry enabled the cutting inserts to
present a strong cutting edge to these stainless steel
testpieces, enabling an undistorted cut path and
ma-chined cusp to be generated This insert geometry
feature, allowed the milled surface topography to be
unaffected by insert inclination angles A stringent test
for any cutter is to machine stainless steel by UHSM
(ie being least × 10 faster than any work previously
undertaken) and here, tests were conducted on
vari-ous grades (Fig 236) The subsequent milled surface
analysis showed little in the way of sub-surface
plas-tic deformation – after UHSM The surface layers
exhibiting only marginal increases in the vicinity of
the surface when tapered sections were micro-Knoop
‘foot-printed’ , over these stainless steel’s substrate (i.e
see Footnotes: 15, Chapter 1; 85, Chapter 7; and Fig
187c – concerning Knoop indentors usage) Milled
surface topography can be viewed visually and surface
parameters taken by either: 3-Dimensional contact; or
non-contact instruments; but in this case, by utilising
an SEM0 with its unique ‘Stereo-imaging and
topog-0 ‘Scanning Electron Microscopes’ (SEM’s), operational
princi-ple is relatively simprinci-ple In that, at the top of the SEM’s column
an electron gun resides having a tungsten filament held in a
strong electrical field This results in the electron gun emitting
electrons (i.e negatively-charged atomic particles), which
ac-celerate to very high speeds These high speed electrons – held
in a vacuum – travel down the column, being influenced by
lenses lower in the column, which squeeze them together to
form an electron beam of very small diameter This minute
diameter electron beam is then focussed prior to colliding
with the test specimen in the microscope’s specimen chamber,
now as a diminutive spot This minuscule spot will then scan
both to the left and to right as well as up and down over the
test surface, the information from which is then brought to a
screen as an image Prior to this, as the electron beam strikes
the test sample’s surface, many different processes occur, such
as: secondary electrons; backscattered electrons; Auger
elec-trons; X-rays; Cathodo-luminescence; etc.; these being
emit-ted, collected and counemit-ted, then utilised for further analyses
raphy software’ – for 3-D visual assessment coupled to its height-to-depth profiling application
For the milled testpieces produced from 316-aus-tenitic stainless steel, subjected to UHSM by this face-mill at 3,000 m min–1, the surface topography showed the influence of the wear land flat produced by the four φ12 mm TiN-coated inserts, although the remain-der of periodic surface offered little sign of any surface modification The 303 stainless steel grade testpieces, indicating a slight improvement over the former 316 grade While, 416 martensitic stainless steel testpieces under identical cutting data generated no appreciable surface blemishes, with the additional benefits of: an extended cutting insert life; significant reductions in both cutting forces and power requirements, over the
303 and 316 stainless steel grades
9.6.2 Ultra-High Speed:
Turning Operations
Introduction
As has been shown in the previous sections with ref-erence to HSM by milling, considerable applied and fundamental research effort has occurred, conversely, little endeavour has been made regarding high-speed turning operations Possibly the major reason for the lack of interest here into HSM by turning operations,
is because a different approach to the workholding is-sues needs to be taken In that, on a CNC turning
cen-tre, or lathe, the ‘bursting-pressures’1 resulting from
significant centrifugal forces with conventional
work-NB The depth of fields from an SEM are considerably deeper
than that produced by conventional microscopes, allowing some exacting surface topography analysis to be undertaken (Source: Smith et al., 2002)
‘Bursting-pressure’ problems, have been well-known in
tra-ditional turning activities for many years This aspect of safe-working practice, was particularly relevant for large cast iron face-plate work, where a rotational speed limitation is imposed by a machine tool builder If this restricted speed is exceeded, then the cast iron – being poor in terms of tensile strength, will literally fragment (i.e ‘burst’), due to the exces-sive centrifugal forces imposed While, the problem is not as severe for chuck and collet work, the lack of gripping-pressure
on the part – at high rotations, will affect the workpiece if long slender parts supported on one end only are turned – possibly causing a ‘whipping-effect’ and attendant safety hazard.