Surface Integrity Cutting Fluids Machining and Monitoring Strategies_4 pdf

9.4.2 Toolholder Design and Spindle Taper Introduction In the past, the taper cone and its associated driving dogs and pull-stud, provided adequate location and torque for the cutter a

Figure 229 Thermal expansion tooling its operation and high-speed turning chuck details

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gal loss of the jaws – when rotating at typically high

turning speeds (i.e see Fig 229b for a diagrammatic

cutaway assembly of an HSM quick-change chuck)

These quick-change chucks incorporate a traditional

wedge-style and lever mechanism, that instead of

dir-ectly acting on the jaws, the radial force acts through

the actuator and lever mechanism, prior to

transfer-ring the effort to the jaws So, when the chuck rotates

at high-speed, the actuators are thrown outward by

centrifugal force, but are restrained from moving by

the lever, which pivots about the central connected

sphere At the opposite end of the lever, the jaws are

also thrown outward and act to move the pivot in the

opposing direction (i.e to that of the actuators) –

ef-fectively balancing each other The performance of the

counter-balanced chuck depends upon the accuracy of

the balance achieved, as the actuator mass is constant

and the top jaw mass being variable depending upon

the particular top jaws in use, thus the state of the

balance will also vary As a consequence, the

clamp-ing force may fall with rotational speed, or actually

increase with heavy and light jaws, respectively With

standard hardtop jaws, the clamping force remains

al-most constant across the range of the operating speed,

making it unnecessary to calculate the clamping force

losses An additional feature is that the static clamping

force can be much lower, since there is no centrifugal

loss This lower static clamping force application, has

the benefit that when turning either thin-walled, or

more delicate workpieces that may otherwise distort

with higher clamping forces, such chucks are unlikely

to affect these components, when an HSM turning

strategy is utilised

Much more could be said concerning HSM

turn-ing operations, particularly relatturn-ing to the calculations

and working practices, but it was not the intention

here, to give a comprehensive account of such

tech-nical aspects, simply a concise account of the

antici-pated problems and possible solutions when turning

at high rotational speeds In the following section, a

discussion concerning toolholder coupling to the

ma-chine tool’s spindle will be briefly reviewed

9.4.2 Toolholder Design

and Spindle Taper

Introduction

In the past, the taper cone and its associated driving

dogs and pull-stud, provided adequate location and

torque for the cutter assembly when mounted into the machine tool’s spindle The tool’s cone taper an-gle was adequately manufactured so that it perfectly

‘wedged’ into its mating spindle taper and the prob-lem of the single-contact mechanical interface was not really exposed as deficient, until very high rotational speeds were being utilised, coupled to much greater feedrates that the newly-developed tooling geometries and tool materials could now exploit In recent years, both dual- and triple-contact tooling systems have been introduced, these designs will now be briefly re-viewed

Dual-Contact Tool/Spindle Design

One of the most significant developments in maintain-ing a complete mechanical interface between the tool-holder and the machine’s spindle was the dual-contact 7/24 taper system The CAT Standard incorporates this 7/24 taper, but also allows simultaneous contact

on both the toolholder’s flange and taper, when HSM machining is the requirement By achieving this dual-contact, the CAT-shank toolholders minimise any form inherent imbalance at say, 2,000 rev min–1 How-ever, if the cutter assembly is to be rotated at 10,000 rev min–1, the toolholder must cope with a × 25 increase

in centrifugal force, which may compound any unbal-ance present in the tooling assembly Further, if the ro-tational speed is increased still further, into the HSM range, then here, the centrifugal force is × 100 greater and the onset of considerable imbalance may create chattering conditions At such high rotational speeds,

if coolant is utilised in the machining process, the HSM conditions could develop a vortex around the cutting tool, that conventional flood coolant pressures cannot penetrate In these circumstances, possibly the only realistic option is to utilise a through-the-spindle coolant delivery application at pressures of >690 kPa (i.e 1,000 psi), coupled to perhaps, micro-filtration of the coolant with special pipes and couplings The CAT system of dual-contact offers reasonable rotational control of the tooling assembly at moderate-to-high rotational speeds, as the mechanical interface system

of face-and-cone provides a certain security against

 ‘Dual-contact 7/24 taper system’ , refers to the taper being to

the 7 inches of taper per 24 inches of length This 7/24 system incorporates several Standards: CAT and BT 40- and 50-taper tooling.

the onset of imbalance Typical applications for these

HSM dual-contact systems include: aerospace part

production; precision die and mould making;

automo-tive component production; as well as medical

compo-nent manufacturing

It is worth digressing somewhat, to explain the

situ-ation of why the single-cone mechanical interface is

simply not effective for HSM production applications

When rotational speeds begin to approach 20,000 rev

min–1, it is not an unusual occurrence for the

single-contact conventional, or standard CAT V-flange

tool-ing assembly to be effectively sucked into the spindle

(i.e as there is no mechanical contact at the flange),

this being the result of a combination of the pull-stud

pressure and the machine’s spindle ‘taper swelling’ – due

to the very high centrifugal force acting at such high

rotational speeds In fact, this minute amount of ‘taper

swelling’ can cause the tool holder to separate from

the spindle’s surface and as a result cause considerable

damage to both the cone’s male and female surfaces

In order to alleviate this HSM problem and run the

tooling assemblies at even faster rotational speeds, the

HSK dual-contact toolholders were developed, which

will now be briefly mentioned

Hsk Dual-Contact Tooling

There are a number of toolholder designs that are

al-ternatives to the conventional steep-taper spindle

con-nection Probably the most popular version for HSM

is the HSK-designed tooling connection (i.e see

perti-nent HSK tooling details in Fig 126c) HSK toolholder

connections offer simultaneous fitment on both the

taper and face, at the front of the spindle The reason

for their acknowledged popularity amongst the HSM

machining companies, is because the increased

rigid-ity of the joint, coupled with their inherent reduction

in dimensions, compared to the equivalent

conven-tional steep-taper connection In Fig 126c, the HSK

8° (included angle) short taper with its gauge face

con-tact and simultaneous taper interference can be seen,

which was designed in Germany to Standard: DIN

69893, being introduced in 1993 HSK is a German

acronym that translates into English as: ‘Hollow short

taper’ Thus, the HSK connection provides:

• both high static and dynamic stiffness,

• offering great axial and radial repeatable accuracy,

• with low mass and stroke,

• having inner clamping

Therefore, with all these proven design advantages over conventional spindle connections, it allows the HSK tooling assemblies to utilise the increased rota-tional speeds necessary for an HSM strategy

Triple-Contact Tool/Spindle Design

The triple-contact connection is being offered by a few toolholder manufacturers (i.e shown in Fig 230) The triple-contact design relies on an inner expand-ing sleeve which maintains uniform contact between the machine tool spindle and the: toolholder’s top ta-per; bottom tata-per; and flange; this being regardless of the spindle speed employed Of particular note is the inner expanding sleeve which functions particularly well at high spindle speeds So, as the centrifugal forces increase – with higher rotational speeds, it causes the spindle to grow (i.e ‘swell’), the toolholder’s spring mechanism forces the split-cone sleeve to proportion-ally-expand with the spindle Further, the expanding sleeve also acts as a vibration-dampening device The expanding sleeve extends the tool’s life on average by between 300 to 500%, by virtually eliminating vibra-tion As a result of this ‘vibration-free interface’ be-tween the tool and workpiece, it provides smoother machining of: tool steels; aluminium alloys; plus other metallic alloys This triple-contact connection system, also performs efficiently with extra-long tools (i.e see Fig 231), notably when utilised on horizontal machin-ing centres The main reason for the enhanced triple-contact tool’s cutting performance with extended tooling assemblies, is the result of the ‘floating’ inner sleeve (Fig 230) which acts to minimise any potential Z-axis deflection, thus maintaining its rotational con-centricity

Such triple-contact tooling is not inexpensive to purchase, but these toolholders really do amortise their cost, by significantly extending cutter life, while improving part production rates Further, it is claimed

by the tooling manufacturer that the toolholder is

‘maintenance-free’ , while its spring-mechanism in

‘life-testing’ has achieved upward of one million tool changes With the advent of either the double- and triple-contact systems, enabling contact between the machine tool’s spindle and the toolholder’s mechanical interface: top-taper; bottom-taper; plus flange; while

‘eliminating vibration’; this has been achieved under the unique conditions that arise with today’s HSM and high-accuracy and precision manufacturing needs

9.5 Dynamic Balance of

Toolholding Assemblies

Introduction

Balancing tools that are intended for HSM

applica-tions is vitally important and there are quite a few

In-dustrial/Manufacturing engineers and users who do not really understand the concept of how to achieve balanced tooling, or why it is really necessary Either very long extended tooling required for say, for deep-pocketing (Fig 231), or tooling that is out-of-balance, will more than likely produce: chattering effects; goug-ing of a step, or face; loss of workpiece accuracy and precision; not to mention uneven and premature cut-ter wear Whenever a new tooling assembly is destined

Figure 230 Triple-contact tool connection

system is ideal for any potential HSM operations [Courtesy of Heartech Precision Inc (HPI)]

Figure 231 Tool runout (≥10 µm) should be of prime importance when machining deep pockets [Courtesy of

Sandvik Coromant]

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for HSM applications on a workpiece, a balancing

operation needs to be undertaken, this statement is

also true for many sub-HSM applications, particularly

when extended tooling is used for whatever reason

(Fig 231) In fact, every rotating object (i.e chuck, or

tooling assembly, etc.), will generate vibration

As has been explained in the previous section, this

vibration results from a number of sources, but

princi-pally here, from centrifugal forces produced by the

ro-tation of an unbalanced mass There are several types

of unbalance that could arise, but here, we are mainly

concerned with what is termed dynamic unbalance,

which increases by the square of the rotational

vel-ocity For example, any vibration produced by a

tool-ing assembly at 3,000 rev min–1, is × 100 greater than

an identical tooling configuration that is rotating at

300 rev min–1 Moreover, what is often either

misun-derstood, or indeed overlooked, is that any change to

the tooling assembly – no matter how small it might

seem, requires re-balancing! These tooling

modifica-tions include any occasion when a cutting tool is

ad-justed, or changed, or similarly if the toolholder is also

either adjusted, or changed Such changes to the

‘sta-tus quo’ of the tooling, will directly affect its ensuing

balance, even minutely when just a ‘few microns’! So

that, these miniscule changes to the tooling’s dynamic

condition, causes a degree of tooling oscillation, hence

an out-of-balance condition – with the likely problems

that this creates

With the wide variety of tooling that is held in: tool

storage carousels; magazines; turrets; etc.; they must

all be ‘balanceable’ by some means A range of

balanc-ing techniques can be employed here for either sbalanc-ingle-,

or dual-plane balancing – more will be said concerning

these effects will be made in the following section The

techniques utilised in achieving tool balance could

in-clude:

• ‘Hard-balancing’ (i.e see Fig 234b) – when the

complete assembly either has to have material

re-moved, or added at a certain part of its assembly

NB The major problem associated with

‘hard-bal-ancing’ is that if the tooling setup changes, so will

the likely rotating mass change, which will mean

modifying the amount of material to be either

added, or subtracted from this newly-distributed

mass,

• ‘Adjustable balancing rings’ (i.e see Fig 232) – by

rotating the twin lower and higher balance rings

either clockwise, or anti-clockwise they minutely modify the balance-condition, allowing single-plane balance to be achieved

NB These matched pair of balance rings are in a

symmetrical state of unbalance (i.e they are both

‘unbalanced’ to the same degree) Letting the user adjust the pair to counter any unbalance in the cut-ting tool/toolholder assembly and locking them into place – usually achieved on commercially-available balancing machines (i.e see Fig 234a) The state of unbalance is not merely a subject to the

‘caprice’ of the machine tool operator, a tool assembly’s balance is given by various quality Standards, such

as ISO 1940/1, or ANSI S2.19 – being basically exact reflections of each other In the following related sec-tions, they deal with how and in what manner rotat-ing cutter assembly balance is achieved, utilisrotat-ing such HSM balance calculations and associated graphical details as necessary, from these Standards

9.5.1 HSM – Problem of Tool Balance

Unbalance of a rotating body (i.e here we are con-cerned with a complete tooling assembly), can be

defined as: ‘The condition existing when the principal mass – axis of inertia – does not coincide with its ro-tational axis’ (i.e shown schematically in Fig 232)

For example, such an undesirable state of affairs can

be comprehended by considering the following situ-ation: if a φ50 mm face mill assembly is rotated at 15,000 rev min–1, it will produce a peripheral speed

>240 km hr–1, which may prove to be disastrous if it is unbalanced!

Basically there exists, three types of unbalance con-ditions for rotating assemblies – such as tooling, these are:

1 ‘Static unbalance’ – single-plane This type of

un-balance occurs when the mass does not coincide with the rotational axis, but is parallel to it and the force created by such unbalancing, is equal to the magnitude at both ends of the rotating body Thus,

if some relief – metal removal (i.e see Fig 234b) –

on the toolholder body equal to the out-of-balance mass that occurs, then a nominal static unbalance is

achieved,

2 ‘Couple unbalance’ – Under these circumstances,

the cutter assembly – mass axis – does not coincide

Figure 232 The taper fitment against runout/eccentricity for a milling cutter and its associated balanced

tool-holder

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with 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:

F = m × r × ( 2 × π × rpm/60) (iv)

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.

both 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.

Figure 233 A graph to determine high-speed cutter unbalance ‘U’ (ANSI S2.19–1989)

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