Smith G.T. Cutting Tool Technology: Industrial Handbook

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face – as it moves to various points on the workpiece’s

geometry. e tooling is held here and in eect, is a

cantilevered beam (i.e. the tool is normally only sup-

ported at one end only). In this situation the tooling

assembly can be equated to that of boring tooling, then

its inherent rigidity will decrease by the cube of the

stand-o distance from the ‘gauge line’

. When a ma-

chine tool builder calibrates this orthogonal-designed

machine tool, they will measure along the base of the

column – either optically (i.e. by for example, autocol-

limation

instrumentation), or with laser interferom-

etry, thus feeding-back any potential error sources for

CNC’s dynamic compensation, during execution of

the part program. Nonetheless, the column can either

minutely: twist; bend; or even move; if its temperature

varies by only just a few °C. In fact, it has been widely

reported that up to 70% of the total errors present on a

machine tool are thermally-related.

On conventional machine tools – with slideways,

essentially one does not really know what is occur-

ring in all of the degrees of freedom at any particu-

lar time. Although even if we did know the dynamic

status, unless some form of ‘adaptive compensation

ystem’

was tted to the machine, having some so-

phisticated multiple sensors positioned on and within

critical positions on the machine’s structure, then oth-

erwise, it would be almost impossible to compensate

18 ‘Gauge line’ of the tool, is the distance from where its eective

length is taken, specically when the tool length compensa-

tion distance is needed for the CNC programming details.

Normally, this ‘gauge line’ is datumed from a known point on

the male spindle taper, located in, and standing slightly proud

of its female counterpart. is measurement is either prefer-

ably set in a tool presetting machine, or it can be obtained on

the machine tool by some form of table-mounted touch-trig-

ger cube. Alternatively, the tool’s length can also be accurately

measured by some form of laser-activated instrumentation,

strategically-positioned at a suitable part of the machine tool’s

structure – where it will not interfere with any subsequent

machining operations.

19 ‘Autocollimation’ was the ‘traditional’ technique that was em-

ployed for any form of machine tool calibration. Autocollima-

tion instruments and equipment, utilise highly-evolved ac-

curate and precision optical apparatus to measure alignments

and squarenesses of the machine tool’s axes, which can then

be utilised to compensate for any error sources detected in the

machine tool’s kinematics. [Sources: Taylor Hobson, 1984/

Spectrum Metrology, Leicester (UK)]

‘Adaptive error compensation’ , basically utilises sophisticated

geometric algorithms that allow for compensation of the geo-

metrical elements, via a range of strategically-positioned sen-

sors on orthogonal machine tools. (Source: Ford, 1993)

for such error sources – as they occur. Clearly then, it

is no coincidence that both the ‘Variax’ operates with

six legs and, it also has six degrees of freedom. ese

leg orientations have three pairs of two legs that cross

each other and which are secured, but free to swivel

at either end (i.e. see the simulated renditions in Fig.

223 and partial assembly of an actual ‘Variax’ machine

tool in Fig. 224). With these six legs, all of the degrees

of freedom are controlled, which in turn, eradicates

the usual sources of errors exhibited by conventional

multiple and ‘stacked’ slideway axes. Some of the ma-

or benets obtained from the ‘Variax’s’ design congu-

ration are:

•

Extremely rigid machine tool – coupled to small

mass making it 500% more rigid than a conven-

tional machine, due to the fact that all forces

whether they are in compression, or tension are fed

through the six legs and its associated space frame,

•

Exceedingly fast in operation – as all the ballscrews

must move together, they only have a light mass to

contend with, making it up to 5 times faster than an

orthogonal machine tool,

•

Very high continuous thrust – due to concurrent

and synchronised use of the six ballscrews,

•

High accuracy and precision – this is due to the

machine’s inherent rigidity, coupled to the fact that

all six degrees of freedom are controlled, with laser

feedback through the centre of the ballscrews

•

No supporting structure – the ‘Variax’ is self-con-

tained, so it does not need foundations. e ma-

chine has a three-point location and the weight

of the spindle head is neutralised, as far as the

ballscews are concerned – by the three large gas-

spring supports

(i.e. see Figs. 224 and 225a).

21 ‘Laser-controlled feedback’ , through the centre of the

ballscrews, negates the Abbé oeset error, with any errors in

the legs (i.e. axes) being averaged, rather than ‘stacked-up’ as

is the situation in conventional machine tools. Hence, volu-

metric accuracies are in line with that of a Co-ordinate Meas-

uring Machine CMM – see details in Fig. 223a.

22 ‘Gas-spring supports’ ,

are tted outside the actuator platform

(i.e. see Figs. 224 and 225a), they are strategically positioned

so that they carry actuator platform’s gravity-induced

load. us, these gas-spring support placements and their

operation, means that the actuators have to only overcome

inertia and momentum to move the machining head. is

has the kinematic benet of enabling the machine spindle to

move through space faster than any single actuator changes

its length, this combination allows the ‘Variax’ to move and

accelerate up to ve times faster than a conventional machine

tool.

Machining and Monitoring Strategies 453

NB e machine tool also has large rubberised

compliance mountings, so when it is in operation

t seems to ‘visually-bounce around’ , but this is

something of an optical illusion, as the overall rigid

integrity of the ‘Variax’ is maintained.

Variax/Hexapod – Specification

and Machining Performance

e ‘Variax’ is equipped with a 630 mm square pallet

and can both move and cut at 66 m min

–1

(i.e. see Fig.

225b), accelerating at >1g while providing >3 tonnes

of continuous thrust in the Z-axis. e machine tool is

compliant with Standard: ASME B5.54, obtaining ac-

uracies of <10 µm

, while its rigidity is >175 Nm µm

–1

Pallet change time is 10 seconds, with a tool change

time of 6 seconds (chip-to-chip), having a tool stor-

age capacity of 50 tools (i.e. expandable). e standard

ool spindle cartridge is rated at 22 kW , with spindle

speed ranging from 100 to 16,000 rev min

–1

, having an

angle of tool spindle inclination of up to 25°.

In demonstrations and in-house trials at the man-

ufacturer’s premises, the initial prototype ‘Variax’ ma-

chine was operated and machining components for

three years. On one large aerospace machining appli-

cation of a critical component, the original cycle-time

was 19.3 hours on a conventional machining centre,

ut when this same part was cut on the ‘Variax’ , it took

just 3 hours to complete, with the additional benet of

being both more accurate and precise. Equally, when

a smaller aerospace component – a landing bracket –

was originally machined it took 1.65 hours on the con-

ventional machine tool, but when the same part was

laced on the ‘Variax’ , it took only 0.55 hours cycle-

time to complete.

Figure 224. The Variax/

Hexapod Machining Cen-

tre here seen during nal

assembly at Nottingham

University, equipped with

an up-rated high-fre-

quency spindle: 40,000 rev

min

–1

@ 40 kW power-rat-

ing, with a ski taper.

NB Laser transducers –

giving an accurate and

precise control, can clearly

be seen located on each

of the actuator legs.

[Courtesy of Giddings &

Lewis/The University of

Nottingham (UK)]

454 Chapter 9

Figure 225. Variax/Hexapod oers a unique and rigid design solution, with many production benets. [Cour-

tesy of Giddings & Lewis]

Machining and Monitoring Strategies 455

e capital outlay for a ‘Variax/Hexapod’ machine

tool, costs about the same as a similar specication –

component size capacity, to that of a conventional ve-

axis machining centre, but with the above additional

performance benets.

Robotic Machining

Robotic machining applications have been utilised for

some years, currently up to a thousand such installa-

tions are to be found world-wide. Probably the biggest

user of robotic machining is the aerospace industries,

although the automotive industry is catching up fast.

Most of the current research work into robot machin-

ing is undertaken with an anthropometric type of

robot conguration, usually having either ve, or six

axes (i.e. see Fig. 226).

If a six axis robot is employed (Fig. 226), it gives

several benets over say, a ve-axis machining centre,

with probably the greatest production advantage being

access to the workpiece’s surface features (Fig. 226b).

is extra degree of freedom, allows the alternative

wrist positions to achieve identical tool positional

orientation – regardless of the workpiece contour,

enabling the robot axis the ability to rotate the wrist

about the tool’s axis. Moreover, where speed and ac-

celeration are important, robots normally out-perform

the more traditional machine tool structures, mainly

due to both their low weight and minimal inertial ef-

fects. Within the robot’s programming language, pa-

rameters exist that allow a balance to be made between

robotic arm speed and accuracy. So, when roughing-

out higher speeds can be employed – at the expense

of accuracy, then the parameters can be changed to a

slower speed, but with greater nished machined part

accuracies.

Limitations also occur when using robots for ma-

chining, with perhaps the most obvious one being

their intrinsic lack of rigidity, when compared to that

of virtually any machine tool. If a robot is employed

at conventional spindle speeds the tool’s cutting forces

are simply too great, creating both vibrations and de-

ections in the robotic arm, which badly impacts on

the component’s machined surface texture and di-

mensional accuracy. In order to mitigate against any

cutting force and accuracy deciencies, it is essential

to utilise HSM spindles, in combination with taking

small D

’s, to minimise these eects.

e robot’s kinematic structural complexity also

causes positional problems, due to innate manufactur-

ing tolerance eects to the location of the joints in the

robotic arm. is dimensional and geometric toler-

ancing build-up, means that the mathematical model

used to control individual joint positions in space,

will minutely dier from the actual reality to that of

the joint placements. is positional dierence being

most apparent and exaggerated around the perimeter

of the robot’s working envelope. Oen, it is impossible

to calibrate the mathematical model for these com-

plex inter-related errors, as the overall complexity of

robot’s control algorithms then become such that they

cannot execute the kinematic motions fast enough to

dynamically control the robot at the required axis tra-

jectories.

In calibration trials on a typical six-axis anthropo-

metric robot undertaken by Young (1999) while work-

ing in cartesian co-ordinates for the robot’s positional

accuracy – when assessed with laser interferometer

instrumentation, the results produced linear errors of

± 0

.8 mm in each cartesian co-ordinate (i.e. X-, Y- and

Z-axes) – across the complete working range. e er-

ror curves produced were in fact, symptomatic of the

kinematic structure of these robots. Characteristics

included a decrease in backlash to almost zero – to-

ward the perimeter of the working envelope (i.e. due

to gravitational eects), while a combination of linear

and sinusoidal eects combined to produce the total

error. Even though these robot errors are large with

respect to those found on a machine tool, they should

not deter robotic usage for suitable HSM applications.

e repeatability shown by most robots is usually far

superior to that of its accuracy.

In practice, any multiple axis robotic arm utilised

for sculptured machining applications, needs to have

their axes biased and oset in order to eliminate the

‘paradox’ that might overwhelm them when all axes

are attempting to keep the end-eector (i.e. tool) nor-

mal to the contoured surface (Fig. 226b). For example,

on a three-roll wrist (Fig. 226), the combination of the

kinematic linear and rotational build-up, may result

in instead of one of these axes angularly moving just

1°, it causes it to actually move 359° instead – thereby

scrapping the workpiece in the process! To alleviate

this problem, if the workstation/stand, is oset to one

side and angled, this oset and compound angle will

minimise the axis predicament that can aict the ro-

bot’s subsequent programmed motions. e worksta-

tion position in its ‘known space’ with respect to the

robot axes datums must be known and accurately cali-

brated, thus ensuring that the angled grid-plate for re-

sulting workpiece xturing is both ‘xed and qualied’.

By slowing down the robotic axis trajectories in the -

456 Chapter 9

Figure 226. Robotic high-speed machining with a multi-axis anthropometric robot, equipped with HSM

milling head. [Courtesy of Southampton Solent University/Westwind Air Bearing Ltd./Smith, G.T.]

Machining and Monitoring Strategies 457

nal machining passes over a sculptured workpiece sur-

face, a certain degree of accuracy can be achieved, but

even here, it does not approach that of any moderately

accurate and precise machine tool. Tool centre-point

programming is the preferred option, as when utilising

cartesian co-ordinates, the programmed points repre-

sent the cutter’s end – once the required oset has been

established. is tool centre-point program allows the

programmer to eectively disregard the tool’s length

in any subsequent machining applications. However,

care must be taken with any exible coupling connec-

tions, such as pneumatics, water, or electrical wiring

to the HSM spindle – allowing sucient slack in the

piping (i.e see Fig. 226), thus ensuring that the robot

will not inadvertently de-couple these services, as it at-

tempts to manufacture the complex part.

9.4 HSM – Toolholders/

Chucks

Introduction

Rotational speeds for tooling assemblies subjected to

HSM applications must of necessity be very high, this

can create several problems for any tooling utilised in

such machining activities. Notably, due to the centrif-

ugal force the toolholder could swell and slacken its

gripping force on the tool’s shank. In an extreme situa-

tion due to the application of cutting forces this might

cause the tool to speedily exit from its holder, in so do-

ing either scrapping the workpiece, or become a severe

safety hazard to any operator in attendance at the ma-

chine tool. erefore, both a good t and connection

is an essential requirement of any HSM machining ap-

plication. us, the mechanical interface between the

toolholder and the machine’s spindle, together with

that of the tool’s shank and its respective toolholder

are the prerequisites for a successful HSM application.

9.4.1 Toolshank Design

and Gripping Pressures

Tool and Sleeve – Mechanical Interface

e latter point briey mentioned in the previous

section concerning the tool shank’s tment in the

holder, is an important criterion in any HSM applica-

tion. Nonetheless, of greater interest and note, is the

actual gripping pressure exerted at the toolholder’s

mechanical interface with its mating tool shank. Some

interesting toolholder designs have been attempted to

increase the gripping pressure here, with the level ex-

erted at the free-end of the holder not being too great

a problem, due to elastic compliance of the sleeve in

this region. e notable diculty arises as the shank

is being gripped toward the ange end of this sleeve,

where with most conventional toolholder sleeve de-

signs, the gripping pressure drops-o signicantly at

this locality. At high rotational speed and under the

application of the cutting forces, the cutter will tend

to become unstable and present a distinct ‘wobbling

motion’ due to the lack of gripping pressure and sup-

port here, being exacerbated by the higher imparted

centrifugal forces. In order to try to alleviate this lack

of grip problem, signicant design eort has been

expended over the years to attempt to increase elas-

tic behaviour and hence amplify gripping pressure at

this sleeve’s region. In the graph shown in Fig. 227b,

the elastic behaviour and its associated contraction are

plotted from the tool sleeve’s free-end in linear steps

along the sleeve and toward the ange. In most prior

designs the sleeve contraction on the tool’s shank near

he ange was approximately 50 µm (i.e. shown by plot

‘B’). By introducing a radially-plunged undercut to the

sleeve’s outside diameter at the juncture of the ange

and sleeve (i.e. see Fig. 227a), this creates signicantly

increased elastic behaviour and hence improved grip-

ping pressure at this region of the sleeve, as illustrated

by the plotted relationship of the test results – produc-

ng a sleeve contraction of 145 µm at the ange, as in-

icated by ‘A’ , in Fig. 227b. is vastly improved grip-

ping pressure with the ‘undercut ange’ is of the order

≈150%, when compared to other design techniques,

which restricts any attempt at the tool’s ‘wobbling be-

haviour’ as a result of the improved elastic contraction

on the shank.

In order to obtain a good overall elastic gripping

pressure along the length of the sleeve, multiple hard-

ened needle rollers are positioned – at a slight angular

inclination, both around and along the male shallow-

angled tapered sleeve’s periphery (i.e. items 2 and 4 in

Fig. 227a). As the female tapered sleeve (item 3) is ro-

tated, the caged inclined needle rollers rotate and be-

gin to steadily move up the male tapered sleeve toward

the ange (item 1) and, in so doing, elastically com-

press the sleeve (item 2), which in turn will tighten

on a tool’s shank. As the tapered female sleeve (item

) is fully tightened with its ‘C-spanner’ , it will com-

458 Chapter 9

Figure 227. High-speed milling toolholders (chucks), holding cutter’s shank along its whole gripping length. [Courtesy of Di-

ashowa Tooling]

Machining and Monitoring Strategies 459

press its face against the rubberised contact seal and

the ange (item 1), thereby acting as both a vibration

damper and sealing against particle/debris ingress into

the mechanism. e front face already being sealed

against such potential particle ingress (i.e. see Fig. 228

sectional detail diagram). In order to increase the elas-

tic behaviour of the male tapered sleeve (item 2 – Fig.

227) and improve its gripping pressure here, the end of

the sleeve is partially slit along its length and around

it at eight equally-spaced positions (i.e. see Fig. 228).

is mechanical design solution for the tightening of

the male tapered sleeve, gives possibly the optimum

gripping pressure for such a mechanical interface on

the tool’s parallel shank.

Tool and Sleeve – Hydraulic Designs

For an alternative tool-gripping pressure design con-

cept, the hydraulic toolholder (i.e. see Fig. 228 – top

right – for a section through the tool sleeve), oers an

ideal alternative to the mechanical tool interface previ-

ously mentioned above. Here, the hydraulic toolholder

anufacturer guarantees a 0.005 mm tolerance-in-

roundness (TIR) at 100 mm from the front face, with

<0.0013 mm repeatability for such hydraulic toolhold

ers. e normal contraction for a φ2

5 mm toolholder is

<0.13mm, although the contraction is proportionally

higher for larger diameters, conversely, it will be lower

for smaller diameters. e benets of utilising hydrau-

lic toolholders are that they provide both rigidity and

balance, while holding the tool’s shank axially-straight

along its own centreline. It is also claimed that hydrau-

lic toolholders produce less vibrations resulting in im-

proved machined surface textures, with a possibility of

less chatter. is latter benet it is claimed to be the

result of the hydraulic uid in the toolholder acting as

a natural vibration damper and impact cushion.

Hydraulic toolholders are balanced to ISO/ANSI

tandards, to G2.5 at 15,000 rev min

–1

in practice, but

are claimed to perform successfully in spindles rotat-

ng at 50,000 rev min

–1

. Moreover, other benets are

made that tool presetting is much quicker to achieve

than mechanical designed toolholder assemblies, as it

is stated that due to the hydraulics having an automatic

centring action, only the tool’s length needs to be pre-

set. Further, as there are fewer moving parts, these

toolholders need little maintenance and the sleeve and

piston should have a fatigue life of >100,000 cycles,

prior to servicing, this ‘in-service life’ representing

many years of practical usage.

e disadvantages to hydraulic toolholders are few,

but may prove a signicant obstacle to their introduc-

tion, including the fact that their purchase costs are

higher than their mechanical counterparts, with the

other limitation being a new hydraulic sleeve is re-

quired for dierent tool shank diameters. However,

this latter point can be mitigated against, by conduct-

ing a programme of: rationalisation; consolidation;

and optimisation on the various production require-

ments for toolholder varieties and shank diameters

(i.e. see Chapter 1, Sections 1.1.1 to 1.1.3 for details of

such tool survey).

Tool and Sleeve – Thermal and Cryogenic Designs

Shrink-t systems require specially made tool hold-

ers, being designed for a specic tool shank diameter,

although they can accommodate any style of shank.

Once in-situ the tool in its thermal sleeve behaves in

a very similar way to that of a one-piece tool. e high

gripping pressure coupled to excellent concentric-

ty (i.e. <5 µm), allows these toolholders to increase

speeds and feeds by >20%, when compared to most

of their mechanically-designed counterparts. Due to

greater rotating concentricity of shrink-t toolholders,

there is a better wear pattern developed on the milling

cutter’s teeth, or drill’s lips, etc., which it is claimed,

increases tool life by >30% over conventional tool-

holders. In a similar fashion to that of hydraulic tool-

holders, thermal contraction occurs both around the

periphery and along the whole mechanical interface,

which will automatically centre the tool’s shank within

its mating bore. is complete toolholder-to-tool t-

ment, minimises centrifugal force when operating in

an HSM mode. us, the contraction of the toolholder

rigidly locks the tool in-situ, this gripping pressure is

at least 500% greater than for conventional toolhold-

ers. In fact the pressure exerted here, is even greater

than that of the pull-stud (i.e. retention knob), mean-

ing that in the presence of high forces, the whole as-

sembly would be pulled out of the spindle before the

tool would be released by its mating holder.

e signicant component of a shrink-t toolhold-

ing system is the induction heating unit, as schemati-

cally depicted in Fig. 229a. Typically, the solid-state

and self-contained unit is relatively compact and in

operation to change tools, the user positions the tool

holder in a receptacle built into a shelf at the front of

the unit. An induction coil is located beneath the shelf

and encircles the toolholder’s sleeve (i.e. collar), which

460 Chapter 9

Figure 228. High-speed milling toolholders (chucks) utilising tapered sleeves with needle rollers for elastic deection,

when tightened – to precisely hold cutter’s shank along its whole gripping length. [Courtesy of Diashowa Tooling]

Machining and Monitoring Strategies 461

quickly heats up the holder and in so doing, expands

the bore allowing the tool’s shank to be placed in-situ

in the toolholder whereupon it contracts around the

shank – taking about 7 seconds to initially expand the

ore by ≈5 µm. is thermal expansion gives enough

clearance to allow the speedy withdrawal of an old tool

and its replacement by the new one. e heat is lo-

calised at the tool’s sleeve enabling the user to remove

it wearing protective gloves. Having removed the new

tooling assembly from the induction machine the user

places this tooling in a solid aluminium cooling-block,

which eciently dissipates the heat from the tooling

assembly. us, just a few seconds aer the assembly

has resided in the cooling-block, the bore shrinks

back, but is constrained from reaching its equilibrium

diameter by the slightly larger diameter of the tool’s

shank. is thermal/mechanical restriction creates the

high, but uniform gripping pressure around and along

the tool’s shank which rmly secures it in place. e

entire shrink-tting process from start-to-nish, only

takes about 30 seconds to complete.

e heating control can be adjusted by the user

to accommodate large, or small diameter toolhold-

ers, requiring diering thermal expansion cycles and

with experience, the optimum settings can be quickly

established. So, once the correct power and tempera-

ture cycles have been created these parameters can

be employed to regulate the induction machine, thus

avoiding energy wastage. If even tighter control is de-

manded of the heating process, an infra-red sensing

device can detect the amount of heat in the toolholder,

so that once it reaches a preset temperature, the in-

duction heating is automatically shut o. Many other

features and protection devices are incorporated into

such induction machines allowing them to be speci-

cally-tailored to the company’s requirements.

Typical HSM shrink-t tool holding assemblies can

e dynamically balanced to G2.5 at 20,000 rev min

–1

while special-purpose toolholders can be purchased

hat are balanced to G1.0 t 40,000 rev min

–1

. More will

be mentioned concerning this critical factor of dy-

namic tool balancing for HSM applications shortly, in

Section 9.5.

Cryogenic methods (i.e not shown) are the direct

opposite of thermal techniques. Here, instead of the

toolholder’s sleeve being thermally-expanded, the tool

is now the subject. As the tool’s shank is inverted and

held at the desired depth beneath the cryogenic liquid

which is held in a suitable container, it is cooled down

by the required ratio of a mixture of liquid nitrogen-

to-methanol. is immersion of the tool’s shank in the

cryogenic uid, shrinks the shank’s diameter. us, the

contracted tool’s diameter is then placed into the tool-

holder (i.e. held at its ambient temperature), which has

its bore as an interference t

with respect to the tool

shank’s diameter. Here, the tool shank nally expands

and causes an increased gripping pressure as it is re-

stricted by the toolholder’s smaller bore. As a safety

note, the user should always wear cryogenically-pro-

tective gloves, otherwise a serious case of prospective

‘frost-bite’ will normally be the likely outcome!

HSM Chucks – for Turning Operations

Although in this section the principal concern has

been with the design concepts and techniques of se-

curely and accurately locking the tool in its respective

toolholder. It is worth deviating somewhat here, to

consider the major problems encountered when at-

tempting an HSM strategy in turning operations, prior

to discussing how and by what mechanical methods,

the rotating toolholders – previously alluded to, are

held in the spindle’s taper.

In HSM turning operations, the grip, or clamping

force exerted by a chuck on the workpiece changes

when the chuck is rotated at speed. is change be-

ing the direct consequence of centrifugal force that acts

on the chuck’s jaw assembly. e magnitude of this

clamping force is very signicant, as a standard power

chuck loses 66% of its ‘static clamping force’ (i.e. the

force exerted by the chuck when it is not rotating) at

its maximum rated speed (i.e. these maximum oper-

ating chuck speeds are dened in European Standard:

EN1550).

23 ‘Interference ts’*, can be best dened, relating to the ‘Limits

and Fits’ of a mating hole and sha, as follows: ‘e minimum

permitted diameter of the sha is larger than the maximum

allowable diameter of the hole.’ (Source: Galyer and Shotbolt,

1990)* In practice, interference ts are those for which, prior

to assembly, the inside (male) component being larger than

the outside (female) component, requires either some form of:

deformation (i.e. pressure); thermal contraction – for a sha;

or otherwise, thermal expansion – for the hole; to obtain a

‘permanent and secure tment’. us, either plastic deform-

ation – for press-ts, or elastic pressure – by thermal eects,

have been exerted between these mechanically-mating sur-

faces, in order to achieve the desired component assembly.

(Sources: BS 4500A and B; Childs, 2004; Griths et al., 2003)

462 Chapter 9