Note: Descriptions are shown in the official language in which they were submitted.
~2~
This invention is in the field of pipelines, and more
particularly concerns a novel mechanical joint for connecting ends
of pipe in telescopingl~ overlapped, frictionally engaged relation.
BACKGROUND OF THE INVENTION
.. . .. .. .
It is known to mechanically connect one pipe with another
by receiving the ends within a concentric mating tubular member
which has an inner diameter slightly smaller than the outer diameter
of the pipe end so that axial engagement of the members requires
large axially-directed force to fit them together. One prior art
method involves carefully preparing the pipe end and the inner
surface of the mating tubular member so that each is a true cylindric
surface and the surfaces are so related in diameter that neither
member is stressed beyond the limit of proportionality. Such form
of joint does not, however, allow use of levated fluid pressures,
and the precision required in shaping the overlapping sur~aces
entails high costs.
For example, a pipe of 100 mm O.D. may be engaged by a
tubular member such as another pipe end or a sleeve, whose I.D. is
not smaller than 99.9 mm in order that the elastic limit is not
exceeded when the members are assembled by being axially forced
together. Ohviously the members may not be out of round and the
surfaces must be of constant diameter, both of which restrictions
make impossible the use of commercial pipe products conforming to
specifications allowing predetermined deviation in diameter, which
may be 0.7 mm.
Another prior art method has been taught in British Patent
Specification No. 435,700 wherein W. Albert discloses assemblv of
-- 1 --
two tubular members ~hose engaging surfaces have diameters such
that one or both of the members is permanently deformed and the
outer dlameter of the overlapping member is equal to the sum of
the outer diameter of the overlapped member plus twice the wall
thickness of the outer member. The strain developed in the deformed
member comprises elastic strain and plastic strain, the stress
corresponding to the elastic strain component causing radial load
between the members, hence the joint is capable of strongly resisting
being pulled apart. To facilitate axial engagement at the start of
the assembly, ~he leading end of either member may be conically
bevelled, or a bell mouth may be formed in the end of the outer
member.
Various implementations of the method above described have
been published, such as, Canada patent NQ . 835,80~ of 1970 to
W.W. Mount, British Patent Specification No. 596,135 (W. C. Youngman),
~anada patent No. 824~865 to C. J. Coberly ahd F. B. Brown, 1969,
Canada patent No. 745,930 to Coberly and Brown, 1966, United States
patent No. 4,328,983 to J. E. Gibson, 1982, and United States patent
No. 4,120,083 to M.C. Echols, 1978.
Certain of the prior art publications require that an end
of a low-carbon steel pipe in ductile state be pre-formed as a bell,
the step producing a permanent set which may e~ceed 6% of the pipe
diameter, and a pipe end is engaged by driving the bell o~er it,
further deforming the outer member plastically. I~ the patent to
Echols it is necessary to produce an inwardly-sloping pipe end to
enable the members to be axially fitted, while in the patent to
Mount the end of the outer member is flared for this purpose.
-- 2
67~
TECHNOLOGY OF THE ASSEMBLY OF TUBES WITH PLASTIC DEFORMATION.
_ . . ~ . . _ _ . . . _ _ _
Prior art joints made by forcing a tubular outer member
over a pipe end so as to distend the outer member have usually
developed permanent enlargement in the range from about 1% to about
7~ or higher, depending on the duc-tili-ty of the metal used. For
example, when an outer member having an I.D. of 96 mm is forced
into o~erlapped relation upon an end of a pipe of 100 mm O. D. the
approximate strain developed, expressed as a percentage of original
diameter, is found to be 4.167%. If a typical low-carbon steel
pipe is used as the outer member, it is subjected to circumferential
deformation comprising both plastic strain and elastic strain, the
- former being much larger. The circumferential tensile unit stress
developed in the outer wall, hereinafter abbrevia-ted as CTUS, will
exceed nominal yield stress of 207 MPa, and will dPpend on the
particular batch of steel, and on the previous mechanical working
during manufacture and final sizing operations. It may be found, for
example, from a stress/strain diagram obtained by elongating a
specimen of the stock, that at a strain of 4.167% the CTUS is
410 MPa, which for a modulus of elasticidy "~" of 207,300 MPa
indicates that the elastic strain component is 0.198%. If the outer
member were allowed to relax to zero CTUS, the permanent deformation
observed would be (4.167 - 0.198) or 3.969~ and the I.D. 99~81 -mm.
The overlapping, highly-stressed outer member subjects the
inner pipe to radial load. If the outer member's wall is 5 mm thick,
this radial force is 41 N/mm2 of pipe surface area, and sets up a
compression stress circumferentially. If the pipe wall is also 5 mm
thick, the unit compressive stress would be greater than 43 N/mm
~. 2~
because its mean diameter is 95 mm. It has been observed that short
concentrically engaged portions of a pipe have a surprisingly great
resistance to deformation, and if the pipe stock is a typical pipeline
steel ~ube manufactured by longitudinal or spiral welding of skelp,
Eollowed by extensive rolling to size the final O.D., such pipe portion
may support the radial load in apparently elastic strain only. The
support provided ~y the non-overlapped pipe beyond the joint contributes
to its resistance to deformation, although the end of the pipe may
be slightly reduced permanently in diameter.
The pull-apart strength of the joint, assuming the over-
lapped length to be 100 mm, and taking a coefficient of friction to
be 0.2, would be (3.1416 x 100 x 100 x 41 x 0.2) = 257,610 Newtons.
A pull of 328,000 N would cause the pipe remote from the joint to
yield, hence the strength of the joint is about the same as if the
pipe ends were welded. If a higher coefficient of friction obtains,
the pull-apart strength would be superior to welded connection.
Despite wide use of pipe joints of the mechanical, friction-
locking type abo~e described, it has not been recognized heretofore
that an internal load exerted on the assembly by fluid pressure in
the pipeline may destroy the joint, even though the pressure may be
well below that which~wo~tld develop yield stress in the pipe wall
remote from the joint, as will be made evi~ent in the following.
EFFECT OF VARYING PRESSURE ON M~CHANICAL JOINTS OF FRICTION-LOCX TYPE
If a joint has been made connecting a low-carbon steel
sleeve with a pipe of similar material by forcing the sleeve axially
over a ring die edge represented by the pipe end, so as to cause
enlargement greater than about one percent of diameter, the CTUS
will range between about 220 MPa and about 700 MPa, while the elastic
strain range will range from about 0.13~ to 0.34%. If the strained
outer member has its stress relaxed to zero while unit stress values
observed are plotted against strain, the curve traced will follow an
approximately straight line, with a slope parallel to the straight-
line portion observed for the metal in its initial ductile state.
Only when the C~US is again increased from zero and raised above the
peak stress value reached during a preceding drawing step will the
curve again follow the pattern of relatively large strain increase
for small unit stress increase, as is typical of ductile deformation.
If the sleeve diameter remains a~ its maximum distention
diameter following assembly, and a fluid is introduced into the pipe
and the fluid pressure is raised, the load created by such pressure
on the joint is partly transferred to the highly-stressed sleeve wall.
If the remanent stress in the sleeve wall is very close to the peak
CTUS reached during the drawing operation, a quite small internal
pressure of the fluid in the pipe will inherently deform the sleeve
further as a ductile member. Even if the elastic strain in the
sleeve has been relaxed slightly so that the CTUS is below its peak
value, the first time that high fluid pressure is applied, as is
required to verify the fluid-tightness of all joints, and which
pressure may develop nearly the yield stress in pipe walls remote from
the joint, the sleeves will be severely deformed and the line destroyed.
It may therefore be understood that for stability in a joint
produced by drawing a sleeve over a pipe end, the assembly must
behave as a system in which a tension spring - represented by the
sleeve member operating within its elastic strain range - is opposed
-- 5 --
by a compression spring - represented by the pipe end engaged by the
overlapping sleeve - also operating within its elastic range, and
wherein the imposition of a load on the system by pressurizing the
pipe does not cause elther member to exceed the peak circumferential
stress value reached during the drawing step. This requirement
clearly cannot be met in an assembly wherein, following the drawing
step, a significant portion of the metal of the sleeve wall remains .
at or near its peak CTUS value.
The fluid pressure which would stress a pipe wall having O.D.
100 mm and I.D. 90 mm to its yield stress, assuming such stress to be
220 M~a, is 24.44 N/mm2. As will be shown at a later point, the joint
defined hereinabove when loaded by fluid pressure in the pipe will
have the pipe wall compressive stress reduced, while the CTUS in the
sleeve wall will be increased, and the load of 2200 Newtons develops
an increase in elastic strain in the sleeve equivalent to its
circumferential elongation by 0.277 mm, or 0.0928~ The new CTUS
under pressure load must not rise above the peak CTUS value of 410 MPa
in the illustration above, hence the state of the assembly when under
no pressure load fixes the CTUS at
0.198 - 0.0982 x 410 MPa = 206.66 MPa.
T.he e~ct of ~he added load on the pipe wall is to decrease its
compressive unit stress; consequently, both members would still behave
truly elastically if the load did not exceed 24.44 N/ 2. However,
in order that the joint be capable of withstanding usual handling
forces, especially bending moments, as well as some torsion, a suffic-
ient safety factor must be included in defining the operating stress
in the sleeve when under no pressure load. Generally, the CTUS value
under no internal pressure should be not greater than about 40% of the
: - '
. . .
peak value reached during the drawing step, hence the objective in
making the assembly herein illustrated would be a CTUS of about 170 ~a
up to about 200 MPa.
The pull-apart strength of the joint, in the absence of
internal pressure load, can reach full pipe yield strength merely by
providing a sufficient engaged length of joint; for example, with
a friction coef~icient of 0.4 under radial load of 17 N/ 2 imposed
mm
on a pipe overlap length of 160 mm; the engaged length can be
reduced because the radial force may rise to above 40 N/mm2 with high
internal pressure, hence an engaged length roughly equal to the pipe
O.D. would suf~ice.
Failure of prior art pipelines connected by mechanical
friction-locked joints may not arise immediately, because during the
application of testing pressure, the assembly remains water-tight both
by the action of sealant compounds and because the radial force
between the tubular members is augmented by the test load; however,
when the test fluid is removed the elastic strain available in the
outer member will be reduced according to the amount of plastic strain
incurred but re-application of pressure without loading the line
longitudinally may not initiate rupture of the seal, and continuously
maintained internal pressure may enable the joints to hold together.
However, under repeated variations of internal pressure, failure of
some or all of the joints is inevitable.
The present invention seeks to improve greatly the reliability
of mechanical, frictionally-locked joints connec~ing tubular members,
and especially to render such joints safe under repeated application
of rated pipe pressures.
~L2~
According to the invention, the assembly of tubular members
by forcing an outer tubular member axially over an end of a pipe
is carried out using a ring die edge formed by the pipe end to
distend the outer member radially by an amount in excess of about 2
of the diameter to fit the outer member tightly overlapped on the
pipe surface, and the circumferential tensile unit stress developed
at peak distention is caused to be relaxed following the assembly
so that the joint will withstand an internal pressure at least as
high as standard testing and operating pressures to which the line
will be subjected.
In carrying the invention into effect in one expression
thereof, the end portion of the pipe at least as long as the length
to bè overlapped is warmed to a temperature above that of the outer
tubular member, after which ~he outer member is drawn axially into
telescoping engagement with the pipe, plastically deforming the outer
member by an amount which slightly but significantly exceeds the
amount that would be produced if the parts were a~ the same temperature.
After a time interval during which the members reach thermal equili~-_
brium, the circumferential tensile unit stress in the outer member
will have decreased below the peak value reached during the drawing
step. Consequently, the outer member is wholly within its elastic
range as a tension spring opposed by the compression spring represen-
ted by the pipe, and the assembly can safely withstand large internal
pressure loading~ and, more importantly, repeatedly varying high and
low fluid pressures. At the same time the radial load between the
members is sufficiently high to develop excellent resistance to
pulling apart.
-- 8
~2~
In an alternative expression of the invention, the heat
produced in the members in distending the outer member permanentl~
and in overcoming the frictional rubbing load that opposes sliding
motion between contiguous surfaces is caused to raise the temperature
of the inner member, on average, more than the average temperature
of the outer member, so that after thermal equalization and return
to ambient temperatures, the circumferential tensile unit stress in
the outer member is suitably reduced below the peak stress reached
during distention.
To this end, a leading end portion of the outer member is
formed with an internal diameter larger than the pipe diameter so
that during the drawing step there is no contact between the sleeve
and the pipe over a length of at least a few wall thicknesses,
causing a strongly-concentrated distention load to be imposed over
a narrow zonet hence effecting strong heating of the pipe surface
along the overlapped length. By chilling the end portion of the
sleeve overlying the zone the temperature difference is enhanced.
When the entire length of the sleeve is also chilled during the
drawing step, the temperature difference at the end of the draw may
be as much as 50Co ~Jhen the leading end of the sleeve is formed with
a conic inner surface, retaining a sufficiently large end face bearing
area to prevent buckling during axial driving, the force required to
initiate distention is lessened as well.
The invention will be more particularly described in the
following explanation of its preEerred embodiments, which is to be
read in conjunction with the accompanying figures of the drawing,
in which:
FIG. 1 is an elevational view in axial diametral section
showing a pipe end and a coupling sleeve positioned prior to assembly;
g _
. -:
7~
FIG. 2 is a diagram showing -the members of FIG. 1 at
initial engagement of a conically-recessed sleeve end with a die edge;
FIG. 3 is a diagram showing the members of FIG. 1 fully
engaged;
FIG. 4 is an enlarged detail section of part of FIG. 3
showing deformation of the leading end of the sleeve;
FIG. 5 is a vector diagram related to FIG. 4 showing the
effect of the overhanging ~art of the sleeve end;
FIG. 6 is a diagram similar to FIG. 3 showing distention
vector distribution during drawing, and showing heating effects
qualitatively;
FIG. 7 is a thermal state diagram of the sleeve wall and
; pipe wall immediately at the end of the drawing step;
FIG. 8 is a diagram similar to FIG. 4 showing an altern-
ative shaping of an end of an outer member prior to drawing,
FIG. 9 is an enlarged section similar -to FIG. 1 showing
a modification o the pipe end;
FIG. 10 is a thermal diagram showing states of sleeve wall
and pipe wall immediately following drawing, when a temperature gradient
has been imposed on the pipe wall; and,
FIG. 11 is a graph relating pressure load effects to the
spring forces of coaxial tubular members in their elastic ranges
Referring to the drawing, a pipe end portion generally
designated 10 and an end portion of a coupling slee~e 11 are shown
aligned and juxtaposed prior to axial assembly. sOth mem~ers are
substantially cylindrical, and may comprise any ductile metal, here
specifically illustrated as low-carbon steel meeting specifications
-- 10 --
67~
for pipeline service. Pipe outer surface 12 has a diameter larger
than ~he_diameter of sleeve inner surface 13, for example by 2~. A
useful xange for the diameter ratio may range from just under 1% to
about 5~ or 6~, or even higher strain on drawing, as will be made
apparent hereinafter. The pipe and sleeve walls 14, 15 respectively
have thicknssses nearly the same; in the present example outer
cylindric surface 16 of sleeve 11 has a constant diameter and is
radially larger than surface 12 by the amount:
~(D - 0.02D - D -~ 2t)
or t - O.OlD
where D is the diameter of pipe surface 12, and
t is the sleeve wall thickness.
The end face 17 of pipe lO is cut along a plane at right
angles to pipe axis 18, and the end face is preferably chamfered by
conic surface l9 coaxial with 1~ and intersecting cylindric outer
face 12, de~ining a circumferential die edge 20.
The left~hand end portion 21 of pipe 11 is internally
chamfered by a conic surface 22 coaxial with axis 18; of the sleeve,
the surface intersecting internal cylindric surface 13 at circumfer-
ential edge 23 and intersecting end face 24 of the sleeve at circ~-
ferential edge 25. The radial extent of the remaining portion of
face 24 is preferably about half of wall thickness "t", although
it may be increased where necessary to withstand application of driving
force. Conic surface 22 is generated on a cone with relatively small
apical angle, i.e. from about ~ to about 22 included angle. The
axial distance between end face 24 and edge 23 should be a small
multiple of -the wall thickness, for example between about 4 and 12
thicknesses, to ensure that the initiation of drawing requires low
driving force, and so that the overhanging leading end portion can
strongly constrict the pipe surface during the drawing step to
improve heat distribution and joint resistance to load by fluid pressure.
In FIG. 2, the tubular members 10 and 11, shown in their
initial state of axial engagement, have been moved together relatively
by suitable means, as by applying a~ially-directed force to opposite
end facP 24' of sleeve 11, while pipe 10 is held in suitable clamps
or pipe slips. The conic surface 22 is shown undergoing plastic
strain by being strongly pressed against die edge 20. Inner edge
23 of the sleeve is slightly distended with respect to the remainder
of sleeve inner surface 13. At the same time the end face 24 is
only very slightly distended, or may even be virtually at its
original diameter. A concavity develops in zone 26 overlying the
pipe end as the sleeve wall i5 radially distended, while narrow gap
27 extends between cone face 22 and surface 12.
In FIG. 3 the members are shown fully engaged, for example
the sleeve extends about 1.3 pipe diameters along the pipe end
portion. A portion of the sleeve approaching die edge 20 is shown
undergoing distention over an axial length of several wall thicknesses
beyond pipe end face 17. The effect of such distention will be
understood from the description following of heating effects produced
at the leading end 21.
The work done in forcing the sleeve over the pipe end
comprises four distinct components producing sensible heat, and these
are reviewed with reference to FIGS. 4, 5 and 6.
- 12 -
~ t7
DEFORMATION_OF LEADING END OF SLEEVE
In the enlarged axial diametral section of FIG. 4 the
outer profile is characterized by deformation of the leading end 21,
shaping it as a circumferentially-extending concavity 26 which
inflects to a convexity 28 that becomes tangent with outer surface
16. Inner sleeve surface 22 is bowed slightly convexly inwardly,
becoming tangent with cylindric inner surface 13 at contact line 29.
The position of line 29 may be slightly nearer -to end face 24 than
is edge 23; the latter becomes essentially obliterated by the
deformation.
The effect of development of CTUS in the sleeve wall, and
particularly in the leading end portion 21, is depicted in FIG. 5
wherein radial load vectors 30.... spa¢ed along pipe surface 12
represent the well-understood loading of an inner cylindric support
body by band tension set up in a contiguous surrounding body. Band
tension is developed also in overhanging sleeve portion 21 extending
to the left of tangency line 29. Unit radial vectors 30'....
progressively increase in magnitude to the right of end face 24,
representing loads acting on a median thickness zone of the sleeve
end prepresented by dashed line 31. At any axial diametral section,
the group of load vectors 30'...represent a cantilever beam loaded
along its length by the circumferential tensions developed, and the
aggregate load transferred to the pipe along a narrow contact band at
line 29 is represented by vector 32. This concentrated load may be
many times greater than the radial load represented by vectors 30,
other than at the die edge 20. Axial movement of the~ leading end
21 of the sleeve therefore pxoduces sharp localized heating of the
pipe surface. The heat generated is shared by the end portion 21,
~ut the effect does not extend axially more than a few wall thicknesses
distance from zone 29 if the drawing step is carried out in 15 seconds
or less time, as is readily feasible.
Referring to FIG. 6, an enlarged scale diagram in axial
diametral section of the assembled sleeve and pipe end is marked
with patterns, each distinctly denoting heat acquisition by one of
four principal modes in which the work of drawing the sleeve over
the pipe is converted to heat. It is to be unde-stood that two
further energy-absorbing phenomena pertain, but are not visually
diagrammable: a minor quantity of work is taken up as stored spring
energy of sleeve wall 15 in circumferential tensile strain, and a
very minor quantity of work is stored in pipe wall 14 in elastic
compressive strain. The patterns are merely illustrative of the
effects and their locations and do not indicate magnitude of heating
other than as suggested by pattern density. Numerical quantities may
be determined by the methods hereinafter presented.
DISTENTION OF SLEEVE WALL
.
Th~ applied axial load forces the sleeve's inner surface 13
radially outwardly as that surface is driven against the narrow die
edge 20; the actual surface length of this edge will quickly decrease
at first due to intense swaging, hut rapid work-hardening, and anti-
welding efects of a coating or applied film 33 tend to stabilise the
zonal width increase and to maintain low friction and wear values.
The sliding motion of the sleeve develops plastic and elastic strain
in wall 15, commencing a few wall thicknesses ahead of die edge 20 and
reaching nearly the peak circllmferential tensile unit stress as the
- 14 -
wall rides over the edge 20. Radial inward deflection of the pipe end,
due to elastic compression, and in some instances due to minor plastic
strain which is observed to be a fraction of one percent, serves to
extend the ramp length over which the wall rides, so that maximum
distension, and peak CTUS, are not reached until the sleeve has
advanced at least about two wall thicknesses beyond die edge 20, at
zone 34.
The drawing step produces therefore an aggregate load due
to the circumferential tension effects at any wall section to the
right of die edge 20, made up of unit vectors 35... and combined
as a single vector 36 acting radially of -the die edge. A sharp local-
ized heating is produced through the region under distention, repres-
ented by oblique dashed lines (legend 37). The heat is readily
computable from the cross-section in tensile stress, the unit stress
value at each section, and the leng~h circumferentially through which
the plastic strain acts. The quantity of heat is however minor, and
at strain values of about 3% it will usually range from 2C to 5C.
FRICTION HEATING WITH DIE EDGE
As set out above, the heat produced by driving the sleeve
over the pipe end is found by estimating a friction factor, and
calculating work as the product with the factor of the circumferential
integral of vectors 36, times the axial length in meters of draw. For
example~ in an engaged length of 120 mm for a pipe edge 20 which is of
diameter 100 mm when 3% strain in the metal develops peak CTUS of 410
MPa, employing a die edge friction factor for treated sleeve surface
at 0.28, the peak radial load if the sleeve wall is 5 mm thick is 41 N
per square millimeter, the unit vector 36 may be 390 N per mm of
- 15 -
~2~61~
die edge circumference, and the work done is 4116 Joules. From the
mass of the sleeve wall receiving a share of the heat, and assuming a
pipe end length of about 35 mm to be heated from the die edge also
sharing the available heat, a wall temperature gain of less than 3 C
results, while the pipe end warms by about 10C over the length given.
The disparity is due to the lesser pipe end mass involved. The actual
temperature of the pipe end may rise at the die edge to nearly 20C,
although heat loss to the relatively cooler sleeve wall will reduce
this; a gradient of temperature along the pipe end will be apparent,
with insignificant heating showing beyond about 40 mm from the die edge.
The shared heat is diagrammed by solid dashed lines (legend 38) which
are sloped reversely to those of legend 37.
FRICTION HEATING OF CONTIGUOUS CYLINDRIC SU~FACES
If as an example the radial unit force exerted by sleeve
surface 13 against pipe surface 12 is 40 N/mm2 and the enyaged length
is 120 mm, the heat quantity developed for a friction factor of 0.4
if a section of axial length one mm is considered in the sleeve wall at
a point just after die edge 20 can be computed theoretically as about
49C, while an equal heat quantity is taken by the pipe surface due to
the movement of this narrow section of sleeve. The actual temperature
attained will be modified because the sleeve surface is continually
encountering relatively cool pipe surface. The last section of the
sleeve to engage the pipe surface will receive n~gli~ible heating.
Similarly, a pipe section just after die edge 20 is subjected to heat
gain during the entire traverse of the sleeve, while at the remote
end beyong line 29 no heat can be gained. The end portion of the pipe
however will be encountering relatively cool new sleeve wall surface,
- 16 -
~L2~6~
hence will no-t reach a temperature increment of ~9C from the rubbing
action. The shared heat of surface rubbing is represented by a
patte~n of short vertical bars (legend 39). The temperature relations
for the pipe and the sleeve are not visually apparent from FIG. 6 but
may be seen from FIG. 7 to be discussed at a later point.
HEATING BY CONSTRICTION ZONE OF SLEEVE END
_ _
The pattern of partial small circles (legend 40) indicates
that a quantity of heat is taken up in the leading end 21 and a
uniformly-distributed heat quantity is taken by the pipe surface 12.
For an assembly with engaged length 120 mm~ pipe O.D. 100 mm and load
vector 32 determined to be 450N aggregate per mm of circumference
acting on the pipe surface with a friction factor of 0.45, the heat
gained in the pipe wall can be computed at about 5.6C while a length
portion of about 40 mm of sleeve end would be warmed by about 20C.
A loss of heat by the leading end 21 is minimized by the gap 27. The
heat gain may moreover be largely offset ~y providing chilling of
outer surface 16 of the sleeve during the drawing step.
HE~T DISTRIBUTION IN ASSEMBLY AT END OF DRAW
_
Referring to FIG. 7, the diagram illustrates es-timated
temperature rise along the sleeve wall and along the pipe wall
produced by the effects set forth hereinabove. In the derivation of
curve 41 showing pipe wall temperature with length, heat quantities
represented by legends 38, 39 and 40 have been combined, with modific-
ation according to proximal surface effects. ~ence the heat gain
concentrated at the die edge has been added to surface rubbing heat,
showing a relatively high sustained temperature for part of the engaged
length. The temperature drops toward the sleeve end hut is elevated
i7~
slightly due to proximity of relati~ely warm sleeve metal.
The heat gain in the sleeve due to wall distention (legend
37) which is a low and constant increment is added to the gains
derived from quantities represented by legends 38, 39 and 40 and is
indicated approximately by curve 42.
It will be evident that the pipe wall temperature increase
is higher than that of the sleeve wall only over a little more than
a half of the length engaged, and that equal temperatures exist at
the crossing point 43. It may be seen when curve 44 is traced wherein
the heat gain due to sleeve end constriction (legend 40), the axial
extent of the superior pipe temperature portion is noticeably and
considerably decreased. Such absence of heat gain would be found in
prior art assemblies lacking the modified leading end form of sleeve,
for example where the end is merely flared to gain entry over the pipe.
Since as will be shown the dimensional changes desired following
a cooling interval must establish a CTUS value in the sleeve wall
below its peak value reached during the drawing step, the possiblity
of effecting an improved degree of dimensional change solely by
bringing the assembly back to ambient temperature from the state
depicted by curve 41 is evident.
The invention may be practiced without using a modified
leading end portion 21 of the sleeve, for example a blunt-ended
sleeve 11 as known from the prior art and shown in FIG. 8 may be pre-
formed to shap~an enlarged entry end with a flared inner diameter,
as shown by curved surface 22' which becomes tangent with inner
sleeve surface 13 at line 23'. The pipe end portion 10 may remain as
previously discussed, or it may be altered as shown at FIG. 9 to gain
- 18 -
7~7
resistance to radial deformation under the larger initial draw force
necessitated by flared surface 22'. In FIG. 9 chamfered face l9 is
considerably extended so that die edge 20 is axially spaced further
from pipe end face 17. The effect of the truncated triangular section
projecting well to the right of die edge 20 is to gain support strength
under the die edge and thereby to minimize radial deflection and to
prevent or minimize plastlc deformation.
As in the embodiment of FIG. 1, a coating or film 33 is
applied on inner sleeve surface 13 prior ~ the drawing step.
Preferred materials include molybdenum disulphide particles adhered
to the metal, or a two-part epoxy resin may be combined to initiate
cure and then rolled or sprayed or brushed on the inner surface, the
material having a sufficiently long cure time so that premature set
does not interfere with the assembly, and preferably requiring several
hours. Such epoxy resin is preferably loaded with a content of very
finely-divided ~iller material in sub micronic particulate form, the
nature of the particles being to counter any tendency of metal surfaces
under high pressures from galling or cold-welding.
When such blunt-ended sleeve is used, the benefit of the
heat gained in the pipe surface according to the embodiments of FIGS.
1 - 3 is not available, nor is the initial engagement as smooth and
requiring minimal axial force. However, as shown in FIG. lO, the
terminal portion of pipe lO may be pre-warmed to set up a desired
temperature gradient as shown by curve 45, the temperature falling
toward die edge 20, while that part of the pipe last to be engaged
by the sleeve is at the highest temperature. This may be chosen in
view of other considerations, but will range from about 12C to 100 C
or more.
-- 19 --
-` ~L2~6'~
Preferably, the die edge and the closely adjacent pipe
material is not heated, or very little.
Any suitable means may be employed to produce the varying
temperature according to curve 45; resistance heaters, steam coils,
or even open flame or heated gas directed on a band of the pipe may be
utilised. As the heated state is highly transient, the use of rapid-
indicating thermometers, such as electronic contact type 'instant'
reading devices should be used. The assembly may then proceed,
and should not take more than about 5 to 15 seconds of time to perform.
It is further advantageous in both embodiments, to reduce
the temperature of the hottest part of the sleeve during the drawing
step; for example, the sleeve end portion 21 may be chilled by being
contacted by a heat-conducting body of any suitable type; or a blast
of chilled air or gas may be directed on the sleeve to lower its
temperature over the end portion thereof.
A curve 142 of sleeve temperature, exhibiting a flattened
left-hand higher temperature portion, and a curve 141 of pipe
temperature which is uniformly higher than the sleeve temperature,
provide excellent dimensional change on equalization of temperature
between the engaged members and return to ambient temperatures, since
the entire engaged length of the sleeve may have a working CTUS set
to a desired value below th~peak stress reached during the drawing
step. This may be understood from following discussion of FIG. 11
- relating to elastic ranges and spring constants of the sleeve and
the pipe wall, and the effects of load due to fluid pressure in the pipe
- 20
In FIG. 11 a straight line 46 originating at the (0,0) point
of an axis system wherein the ordinate is calibrated in Newtons force
and the abscissa is in mm of elastic deformation of an assembly of
a stressed sleeve engaged over a pipe, represents spring load when
both members are in elastic deflection.
At point 47, another straight line 48 intersects line 46,
and line 48 intersects the abscissa at point 49. The line represents
the compressive elastic stxain in a pipe engaged by an elastically-
stressed sleeve. The sense of the deflection of the~compression
spring is to be understood in that, if the spring end were free, it
would lie at point 49 under no spring load, while when it is engaged
by the tension spring, they are tied together in opposed relation under
a common deflection point represented by 47. The loads on both springs
are equal.
Let it be assumed that the condition depicted represents a
CTUS in the sleeve spring at its peak stress value before relaxation.
In other words, the load o~ point 47 is the maximum the sleeve can
sustain, without being plastically strained.
If, after relaxation, the compression spring and tension
20 spring are depicted by line 148 intersecting line 46 at point 147,
and intersecting the abscissa at point 149, the significance is that
their common load is reduced, and that the CTUS in the sleeve is
lowered, while the compressive stress is also lowered. I now a given
load in Newtons is applied to the system, namely by erecting a vertical
line 150 whose ordinate value between its intersections of line 46 and
line 148 is a load due to pressure in the pipe, the upper point
intersects line 46 at 147', denoting a CTTJs just below thepeak stress
initially imposed during drawing. The system is therefore safe from
overloading, regardless of the number of cycles of applying the load,
and the sleeve remains in its elastic state at all times. The
pipe remains at reduced compressive stress.
- 21 -
