Performance of TMCP steel with respect to mechanical properties after cold forming and post-forming heat treatment * David Porter a, * , Anssi Laukkanen b , Pekka Nevasmaa b , Klaus Rahka b , Kim Wallin b a Rautaruukki Oyj, Box 93, FIN-92101 Raahe, Finland b Materials and Structural Integrity, VTT Industrial Systems, P.O. Box 1704, Vuorimiehentie 5, Espoo FIN-02044, VTT, Finland Abstract The paper describes the results of work done in the Finnish part of the ECOPRESS project on the mechanical properties of pressure vessels made from the TMCP steel grade P420ML2. Dished end (DE)-cylinder assemblies with diameters of 2500 mm and thicknesses of 15 and 30 mm have been examined using DE in both the cold-formed (CF) and CF and post-forming heat treated conditions. The blanks for the DEs contained welds to enable the effect of cold forming on weld metal to be evaluated. Cold forming increases both the transition temperature and the strength of the DEs. Nevertheless, toughness against brittle fracture in the CF state is good for all parts of the DEs and girth weld with T 27J !K50 8C in the absence of blank welds and !K20 8C when blank welds are present. For PFHT DEs T 27J !K50 8C even for the blank welds. T 27J was found to correlate with the fracture toughness reference temperature T 0 which can be used to determine the minimum operating temperature. The impact toughness of the CF DE can be determined by compressing plate specimens 15% and ageing 30 min at 250 8C. Such a procedure can form the basis for an additional requirement on PxxxM/ML grades when high cold forming strains are involved in vessel manufacture. Ductile fracture is not of concern, as upper shelf toughness remains high in all parts of the DE. The yield and tensile strengths of a CF DE are much greater than those of the cylinder, whereas the membrane stresses on the cylinder are greater than those on the DE. Consequently, design can be safely based on the properties of the nominally unformed cylinder. Furthermore, tensile instability will be first reached in the cylinder before it is reached in the DE, even though Y/T is at its lowest in the cylinder. Secondary bending stresses are greatest at the DE knuckle, but the CF DE has more than sufficient ductility in bending to accommodate the bending strain. The high Y/T of CF DE is combined with high material ductility and is therefore fully acceptable. q 2004 Elsevier Ltd. All rights reserved. Keywords: Dished ends; Cold forming; Heat treatment 1. Introduction In common with the other parts of the ECOPRESS project, the purpose of this part was to help demonstrate that high-strength grades of steel can be applied in pressure vessels. Here the properties of the thermomecha- nically processed grade P420ML2 (EN10028-5) were examined. The use of such steel should allow higher design stresses than can be reached with conventional normalised grades like P355N, while offering the advantages of superior weldability over the normalised grade P420N. Thermomechanical processing is known by the acro- nyms TM or TMCP, i.e. thermomechanical controlled process. Steels made in this way are characterised by excellent combinations of strength, toughness and weld- ability. The lower carbon equivalent of TMCP steels compared with conventional normalised grades of equiv- alent strength means easier welding with an increased safety against hydrogen cracking, lower preheat, less need for repair welding, etc. Also, the low carbon and impurity contents of TMCP steels mean that they are well suited to 0308-0161/$ - see front matter q 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijpvp.2004.07.006 International Journal of Pressure Vessels and Piping 81 (2004) 867–877 www.elsevier.com/locate/ijpvp * 29th MPA Seminar, Stuttgart, October 9 and 10, 2003-ECOPRESS Seminar. * Corresponding author. Tel.: C358(0)20 59 22 266; fax: C358(0)20 59 23 101. E-mail address: [email protected] (D. Porter).
Transcript
Performance of TMCP steel with respect to mechanical properties
after cold forming and post-forming heat treatment*
David Portera,*, Anssi Laukkanenb, Pekka Nevasmaab, Klaus Rahkab, Kim Wallinb
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877868
cold forming. There are more than two decades of
experience with the application of TMCP steels, which
have been extensively used in linepipe in the oil and gas
industry and as structural steels in, e.g. bridges or
offshore. Unlike in Japan, in Europe TMCP steels are,
as yet, rather uncommon in the pressure vessel industry.
This is largely due to a lack of information and
experience within the industry. One example of their
application is given in Ref. [1].
Thermomechanical processing means the control of
temperatures and rolling reductions during hot rolling and
the control of cooling after rolling. Combinations of
strength and toughness are obtained that are inaccessible
by heat treatment alone. As a consequence, the heat
treatment of TMCP steels is a little more limited than it is
for normalised or quenched and tempered steels. Heat
treatments below the lower critical temperature Ac1, like
post-weld heat treatment, are permissible, however. The
only limitation of practical importance is that the steels
cannot be normalised. This may have led to a misconception
about the use of TMCP steels in pressure vessels, especially
in dished ends (DE).
In the case of older, lower toughness, conventional
steels, it is necessary to normalise the steel after cold
straining beyond 5%, in order to restore the original
toughness and to protect against a further loss of toughness
through strain ageing during subsequent welding. In the
case of DEs equivalent cold strain can reach 25% or more.
Obviously, if TMCP steel behaved in the same way as
conventional steels, they would be unusable in cold-formed
(CF) DEs. However, previous experience [2] has shown
that TMCP steel has such a good toughness reserve that it
can meet low-temperature toughness requirements even
after high degrees of cold strain and the subsequent ageing
caused by welding. As shown below, the results of this
ECOPRESS project demonstrate that TMCP steels can be
CF into DEs and subsequently welded to the PV cylinder
and still meet relevant strength and toughness requirements
even without the need for any heat treatment. Furthermore,
it is shown that, if required, mechanical properties can be
enhanced by the application of post-forming heat treat-
ment. The presentation addresses the central issues of
toughness against brittle and ductile fracture and plastic
instability.
2. Materials and vessel manufacture
Full experimental details are given in Ref. [3] with
only a brief description here. The plates used were
representative of the two main types of TMCP steels: a
15 mm thick plate that was cooled in air after rolling, and
a 30 mm thick plate that, after rolling, was water-cooled
to 580 8C at 8 K/s and then cooled in air. Both were
thermomechanically rolled. Tables 1 and 2 show their
chemical compositions and tensile properties. The plates
were welded together to make 3100 mm diameter round
blanks, which were then dished (Fig. 1) and flanged (Fig.
2) to make 2500 mm diameter, torispherical DEs of the
‘korbbogen’ type. For such ends, the crown radius is
0.8!2500 mm and the knuckle radius is 0.154!2500 mm. For each thickness, two DEs were made: one
was left in the CF state and the other was given a PFHT
at 575 8C with a holding time at temperature of 2 min per
mm thickness. All four DEs were then welded to 300 mm
high cylinders made from the same plates, Fig. 3. The
blank welding and the girth welding of the DEs to the
cylinders was done using submerged arc welding with
OP121TT flux and S2Ni2 wire from Air Liquide. Heat
input ranged from 1.7 to 2.5 kJ/mm with most runs close
to 2.0 kJ/mm. With respect to the base plate properties,
the chosen weld metal is overmatching in the as-welded,
all-weld-metal condition, for which YSZ510 MPa and
UTSZ605 MPa. The manufacture of the DEs, post-
forming heat treatment and welding was done at the
Halikko works of Rautaruukki.
The round blanks were marked with a square grid of points
100 mm apart to allow the determination of overall forming
strains after dishing and flangeing. Measured principal
strains and the calculated von Mises equivalent strain are
shown in Figs. 4 and 5.
Fig. 3. Dished end welded to cylinder.Fig. 1. Dishing.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877 869
3. Evaluation procedures
All four DE-cylinder assemblies were tested using
samples from the crown, knuckle and flange regions of the
DEs together with samples from various zones across
the girth and blank welds. The crown was also tested after
ageing to simulate the effects of fixture welding. In addition,
the base plates were tested as delivered and after numerous
combinations of tensile and compressive straining, ageing
and high-temperature heat treatment. Testing included
hardness traverses, tensile tests at 20 and 300 8C together
with true stress–strain measurements, impact testing with
various sample/notch orientations, determinations of To
temperatures using Charpy-sized fracture mechanics speci-
mens, ductility testing with notched bend samples and
metallographic examinations. Toughness testing was mostly
done using sub-surface specimens transverse to the plate
rolling direction. Both conventional, through-thickness-
notched specimens (TL) and surface-notched specimens
(TS) were used. The TS orientation was considered to better
Fig. 2. Flangeing.
produce results representative of the toughness seen by non-
penetrating surface cracks, while the TL orientation is that
presently used in practice. Full Charpy V transition curves
were established to enable the determination of the 27J
(or 28J) transition temperature and the upper shelf energy.
In the case of the welds, preliminary tests in the transition
regime were used to find the most brittle of the various HAZ
sub-zones (fusion line (FL), coarse-grained HAZ, inter-
critical HAZ). This was subsequently used for full transition
curve determinations. In practice it was either the FL or the
CGHAZ. In addition, full transition curves were determined
for the FLC10 mm position, where the ageing effect from
the heat of welding is greatest [2]. Specimens from the DE
flange were located at about the FLC40 mm position with
respect to the girth welds (GW).
Notched bend testing using a notch radius to net
section ratio of 0.25, was done for the purpose of finding
the limiting local strain criteria for a strain state that is
more damaging than that prevailing in tensile testing,
Fig. 4. Definition of strain measurement.
Fig. 5. Distribution of overall forming strain in the 15 mm thick dished end.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877870
see Fig. 6 below. The test geometry resembles that used
initially by Giovanola et al. [4]. In the present work,
however, a calculation methodology was developed to
allow the conversion of the measured value of total/
elastic deformation in the bend specimen to be
transformed into the value of the ultimate strain at the
notch tip. The latter corresponds to the initiation of
tearing and is ascribed to the point of maximum load in
the bend test. A finite element analysis of the notched
specimen using the measured true stress–strain behaviour
of the test material was used in the development of the
methodology. Detailed procedures and results are given
in the background document to the ECOPRESS project
[3]. The most important results and conclusions are
summarised below.
4. Impact and fracture toughness
Figs. 7 and 8 summarise some of the results from the
Charpy V testing of the base plate and the DE-cylinder
assemblies. Results for TL specimens are shown here as, on
average, these gave T27J transition temperatures that were
8 8C higher than the TS specimens (Fig. 9). The results in
Fig. 7 show that all parts of the DE-cylinder assemblies have
low values of T27J. In the case of the 30 mm thick material
(Fig. 7b), T27J remains remain below K50 8C in all parts of
the assembly even without PFHT. The same is true of the
15 mm thick material with the exception of the weld metal
Fig. 6. Details of the three-point notched bend test conditions. sZ20 mm
for the 30 mm plate and sZ10 mm for the 15 mm plate.
and FL of the cold worked blank welds, Fig. 7a. For these
T27J is below K20 8C when the DE is left in the CF state.
However, PFHT lowers the transition temperatures of the
blank welds to below K50 8C.
Upper shelf toughness remains high in all parts of the
DE, despite the presence of high forming strains and ageing,
Fig. 7. 27J transition temperatures for cold-formed and PFHT dished
end–cylinder assemblies. Transition temperatures are subzero. (a) 15 mm
thick material, (b) 30 mm thick material.
Fig. 8. Upper shelf energies for cold-formed and PFHT dished end–cylinder
assemblies. (a) 15 mm thick material, (b) 30 mm thick material.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877 871
Fig. 8. Consequently, ductile fracture is not expected to be a
cause for concern.
Figs. 7 and 8 show that apart from the effect on the blank
welds, the effect of the applied PFHT is quite small. Details
concerning the effect of different PFHT temperature and
time combinations on toughness and strength can be found
elsewhere [3]. It should be noted that PFHT is done to the
DE before girth welding and it is therefore quite different to
PWHT. The differences in KV (US) and T27J between the
CF and PFHT states for the GW in Figs. 7 and 8 must be
due to statistical variation or differences in the weld bead
location, for example.
Fig. 9. The effect of specimen orientation on 27J transition temperatures.
Statistical Master Curve analysis was made on data from
static three-point bend fracture mechanics specimens for the
base plates, the base plates after various degrees of
homogeneous stretching and ageing, the DEs and the
CGHAZs of the GW. For the DEs, the flange, crown
and aged crown were studied. In the case of the GW, the
CGHAZ of the GW between the cylinder and the PFHT DE
was studied as this had shown the highest values of T27J. As
discussed above, the results should be equally valid for the
GW of the CF DE. TS specimens were used throughout with
the exception of the unstrained base plate, where TL
specimens were also used. T0 fracture toughness reference
temperatures (KICZ100 MPa m) determined using the
ASTM standard procedure [5] were found to be lower
than when applying the SINTAP procedure [6]. This was
especially true in the case of the GW CGHAZ specimens
and presumably results from the microstructural inhom-
ogeneity below the fatigue crack in the region of the
FL/CGHAZ. Consequently, the SINTAP approach for
determining T0 is recommended. The T0 values determined
using the SINTAP procedure are shown in Fig. 10 together
with corresponding values of T27J. On average, the
difference between the two temperatures is 25 8C. This is
close to the mean difference of 188 found for a wide range of
steels and places the TMCP steels in a conservative part of
the scatterband.
Fracture mechanical evaluations using the Master Curve
method have been made to assess the performance of CF
DEs in the ductile to brittle transition region. The existence
of a pre-existing semi-elliptical surface crack was assumed,
with a crack depth of 25% plate thickness and length/depth
ratio of 3. Residual stresses were taken as equal to the yield
strength of the base material. Fig. 11 shows the results of
two such analyses. For an applied stressZRm/2.4 (in
accordance with prEN 13445-3), the results indicate that
even for a conservative cumulative failure probability
Fig. 10. Correlation between the fracture toughness reference temperature
T0 and T27J. Both 15 and 30 mm materials. BP, unstrained 15 and 30 mm
base plate. CF and aged, cold-formed and aged base plate (stretching 0, 10
or 15%; with and without ageing 30 min at 250 8C; both 15 and 30 mm
materials). DE, flange, crown and aged crown in 30 mm material. GW
HAZ, coarse-grained HAZ on the DE side of the girth weld between PFHT
DE and cylinder, 15 and 30 mm materials.
Fig. 11. Effects of design stress on the lowest permissible operational
temperature for case with yield strength 420 MPa. Design stresses,
500 MPa/2.4Z208 MPa and 0.8!420 MPaZ336 MPa.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877872
specification PfZ0.05, and a plate thickness B of 30 mm,
the difference between the 28J transition temperature and
the lowest operational temperature (ToperationalKT28J) is 08.
In other words, Toperational is equivalent to the T28J attained
Fig. 12. The effect of straining and ageing on T27J for subsurface TS specimens,
star-shaped symbol correspond to the given DE locations.
for the base material or welds. For the 15 mm plate, the
corresponding difference is about K50 8C, giving very low
minimum operational temperatures. Increasing the applied
stress to 0.8!yield strength causes the lowest operational
temperature to increase by an almost constant value of
about 10 8C.
As a result of the cold forming (and ageing) the actual
yield stress of the CF DEs is much higher than the specified
minimum value (SMYS). Therefore, the possibility exists
that there will be higher residual stresses than the SMYS in
the structure after welding. Residual stress patterns have not
been studied in this project and it is uncertain to what extent,
for example, GW residual stresses will be governed by the
yield strength of the DE flange. Normally, residual stresses
are assumed to be at most equal to the yield strength of the
weld metal. With the welding consumables used here, the
yield strength of the weld metal could reach 570 MPa at
20 8C. Analyses of higher residual stresses showed that
increasing the residual stress from 420 to 650 MPa, with
failure probabilities of 0.05–0.2, raises the lowest oper-
ational temperature by approximately 15–20 8C, as far as
plate thickness relevant to the present TMCP steels are
(a) 15 mm thick material, (b) 30 mm thick material. Open symbols and the
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877 873
concerned. This means that even in a most severe case with
(i) considerable strengthening as a result of cold-defor-
mation and ageing, (ii) a corresponding elevation in welding
residual stress, and (iii) a 30 mm plate thickness, Toperational
is still only 15–20 8C above T28J for the base material or
welds. Considering that even the highest T28J that was
recorded for the DE GW, was as low as K60 8C, the lowest
Toperational would thereby be K45 to K40 8C. This should
therefore provide a sufficient toughness reserve and low
enough operational temperature for most structural PV
applications.
Fig. 12 shows the effect of straining and ageing on T27J
for the DEs compared to base plates aged 30 min at 250 8C
after straining in tension 2, 10 and 15%, and in compression
10, 15, 20 and 25%. Tensile stretching was applied
transverse to the rolling direction, while compression was
through-thickness plane strain. The data has been plotted as
a function of the von Mises equivalent strain in an attempt to
compare the strain states in the different specimens. The
values for the DE crown are those obtained after ageing
30 min at 250 8C. The results marked ‘DE flange’ are for
specimens 10 and 40 mm from the GW FL that have not
been artificially aged but they are comparable to the
artificially aged strained base plates due to the ageing
caused by girth welding. It is clear from the results that T27J
is not a linear function of strain. The approximately linear
behaviour over the range of strains that can be applied in
tension before necking cannot be extrapolated to larger
strains as is shown by the compression results. Furthermore,
it can be seen that T27J for the various parts of the DEs is in
Table 3
Tensile test results from CF DE and base plate
Test
temp (8C)
Plate
(mm)
Position Direction
Rel RD
State ReH
(MPa)
20 15 Base plate trans as-del. 463
Crown trans as-del.
Crown trans aged 250C
Flange trans as-del.
Girth weldb n.a. as-del. 573
Blank weldc n.a. as-del.
30 Base plate trans as-del. 445
Crown trans as-del.
Crown trans aged 250 8C 516
Flange trans as-del.
Girth weldb n.a. as-del. 521
Blank weldc n.a. as-del.
300 15 Base plate trans as-del.
Crown trans as-del.
Flange trans as-del.
Girth weldb n.a. as-del.
Girth weldc n.a. as-del.
30 Base plate trans as-del.
Crown trans as-del.
Flange trans as-del.
Girth weldb n.a. as-del.
a Y, Rp0.2 or ReL when Rp0.2 not given; T, Rm 10 mm diameter round specimensb Weld specimens extracted parallel to the weld.c Blank weld specimens taken from the knuckle location.
good agreement with T27J given by compression testing to
the same level of equivalent strain. The reason for the
deviation between the tensile and compression results for
the 30 mm material is unknown. Ignoring these, all other
data points fall around the parabolic fits that show a
maximum effect at an equivalent strain of about 0.20, i.e.
close to 15–20% compression strain. This non-linear
behaviour is also in line with previous experience [7]. The
maximum shift of T27J is about 40 8C for the 15 mm material
and about 25 8C for the 30 mm material. The differences in
the shifts may be explained by the lower carbon content of
the thicker water-cooled material (0.07 compared to 0.09%)
together with the use of titanium microalloying in the
thicker material.
The above results imply that the impact toughness of a
CF DE can be determined by compressing plate specimens
15% and ageing 30 min at 250 8C. Such a procedure can
form the basis for an additional requirement on PxxxM/ML
grades when high cold forming strains are involved. When
300 C 15 Crown trans as-del. 366 513 26.7 68.6 0.71
Crown trans aged 250 8C 366 513 27.1 70.6 0.71
Flange trans as-del. 455 558 23.1 77.2 0.82
Girth weldb n.a. as-del. 349 553 28.1 71.1 0.63
Blank weldb n.a. as-del. 480 592 26.5 59.0 0.81
30 Crown trans as-del. 392 516 28.1 79.2 0.76
Crown trans aged 250 8C 394 515 22.5 79.2 0.77
Flange trans as-del. 453 537 22.1 75.0 0.84
Girth weldb n.a. as-del. 455 613 27.5 57.3 0.74
Blank weldb n.a. as-del. 431 440 582 25 62.0 0.76
Blank weldb n.a. as-del. 435 556 17.4 60.5 0.78
a Y, Rp0.2 or ReL when Rp0.2 not given; T, Rm 10 mm diameter round specimens.b Weld specimens extracted parallel to the weld.c Blank weld specimens taken from the knuckle location.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877874
forming strain and values of A5 below the 19% specified for
the base plate are unavoidable. Y/T values are also high after
cold forming, which leads to low uniform elongations, but
nevertheless reductions of area at fracture (Z) remain high,
essentially unaffected by the high degrees of cold strain and
Y/T values. A similar effect was apparent in the case of upper
shelf toughness. The lowest values of Z are found in the weld
metal, especially at elevated temperature. However, these
values are quite normal and are due to the higher oxygen
content of the weld metal compared to the base plate.
The application of PFHT to the DEs shifts the tensile
properties in the direction of the base plate, with
corresponding increases in A5 and reductions in Y/T, see
Table 4 and Fig. 13. Again, Z values are largely unaffected.
Fig. 13. Yield and tensile strengths at 20 8C for all the tested parts of the 15
and 30 mm dished ends (From Tables 3 and 4).
Tables 3 and 4 also show that yield strength drops as the
temperature rises from 20 to 300 8C, but tensile strength
drops much less, if at all. Unstrained base plate shows the
greatest change in yield strength of all the parts of the DE-
cylinder assemblies. Therefore, if the vessel is dimensioned
on the basis of the base plate yield strength at temperature,
the strength margin to yield for the as-CF or PFHT DE
will, in fact, be higher at 300 than 20 8C. The margin
between the design stress and the tensile strength will be
even greater.
Fig. 14. Load-deflection record for notched 3P bend test specimen. Vertical
axis is the primary bending stress sZF/2!4s!6/(3s!s2)Z4F/s2 (F is
load and s is specimen span) and horizontal axis is measured total
deformation/elastic deformation NMOD/NMODelastic.
Fig. 15. Local strain as a function of relative deformation of 3P bend bar,
KtZ1.8.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877 875
Those parts of the DE-cylinder assembly with the
lowest Z values at 20 8C were selected for bend testing.
These were the blank welds in the PFHT DE and the GW
in the 30 mm DE. A typical load vs. notch mouth opening
displacement (NMOD) record is shown in Fig. 14 with
NMOD given as the ratio of the total to elastic
components.
The corresponding calculated principal and equivalent
strains below the notch are shown in Fig. 15. Maximum load
values of NMOD/NMODelastic for the three specimens
studied are given in Table 5 together with corresponding
ultimate notch tip strains expressed as equivalent values of
the reduction of area (Zequiv). From the table, it can be seen
that these Zequiv values range from 66 to 89% of the
corresponding uniaxial reductions of area at fracture. The
effect of the plane strain conditions below the notch is thus
to moderately lower the ultimate strain capacity of the
material relative to that obtained under axisymmetric tensile
conditions. However, the actual ultimate strains achieved
show that a very large deformation capacity exists, as is
apparent from the large limiting deformations in bending.
The limiting bending load section stresses correspond
roughly with the true stresses at the respective strains.
This shows that the bend resistance of a notched section still
increases in accordance with the material’s hardening
capacity up to the ultimate ductility for strains well beyond
the uniform elongation measured in the simple tensile test.
Thus a decreasing load resistance like that that occurs in
tension due to the onset of geometric softening after
Table 5
Results from the notched three-point bend test
Specimen Uniaxial tensile test
ReL Rp0.2 (MPa) Rm (MPa) A5 (MPa) Z (
15 mm blank weld PFHT 497 576 21.5 74
30 mm blank weld PFHT 496 565 22 73
30 mm girth weld (PFHT) 518 565 25 76
the tensile load maximum does not operate in a section
under bending. Consequently, the inherent safety of a CF
knuckle region is a fact.
The strength difference between the DE flange and the
cylinder is reflected in hardness profiles across the GW,
an example of which is shown in Fig. 16. This, and similar
profiles from the other welds, show that the welds are quite
homogeneous, with relatively small hardness differences
between the different zones. On the cylinder side of the
welds, the hardness increases uniformly up to the FL with
no soft zones. On the CF DE side of the weld there are signs
of narrow zones with a locally slightly low hardness due to
the recovery of the cold worked ferrite at temperatures
around AC3. These shallow ‘soft’ zones are still harder than
the cylinder material and have no practical effect on the
strength of the welds. This was demonstrated by cross-weld
tensile tests, which fractured outside the weld zone in the
lower strength cylinder.
One significant element in the safety of a pressure vessel
is the design margin to rupture in the design state, i.e. a
vessel nominally without strength-deteriorating flaws.
Stresses in a cylindrical vessel with DEs vary from an
undisturbed membrane biaxial 2/1 tension in the cylinder to
equibiaxial in the crown of the DE. In the end of the cylinder
the hoop membrane stress becomes lowered for geometrical
reasons and, in the knuckle, the hoop stress may even
become negative if the knuckle radius is small. When using
a CF DE, the usual choice for section thickness obeys the
minimum thickness at the crown and that is set to the same
as the nominal thickness of the cylinder. Design thus obeys
the strength of the nominally unformed cylinder, where the
actual yield strength is the lowest of all parts of the vessel.
The design margin for a cylindrical vessel with a CF DE
may be conservatively considered by relating the membrane
stresses to the yield strengths of the different regions
(Fig. 17). It is easily seen that the smallest margin against
yield occurs in the cylinder, while the elevated yield
strength in the flange and knuckle effectively increases the
design margin for the DE.
The high yield and tensile strengths of the CF DE
compared with the tensile strength of the cylinder means
that tensile instability will be reached in the cylinder before
it is reached in the DE, even though Y/T is at its lowest in
the cylinder. Indeed, burst tests outside ECOPRESS have
demonstrated that CF DEs remain almost in shape while
burst occurs in the cylinder [8].
Notched 3PBT
%) Y/T (%) NMOD/NMODel Zequiv (%) Zequiv/Ztensile
0.86 32 68 0.89
0.88 30 63 0.86
0.92 22 50 0.66
Fig. 16. Hardness profile across the girth weld of the 15 mm thick material.
Fig. 17. Design margin of a cylindrical vessel with dished end produced by
cold forming. Cold forming increases the yield strength in the dished end
which together with the local stresses and stress states causes the design
margin to increase above that obtained for the essentially unformed
cylinder material.
D. Porter et al. / International Journal of Pressure Vessels and Piping 81 (2004) 867–877876
Secondary bending stresses are greatest at the DE
knuckle, but these need not be considered as the CF DE
has more than sufficient ductility in bending to accommo-
date the bending strain.
6. Conclusions
The work reported here has been concerned with the
properties of the TMCP pressure vessel grade P420ML2
after cold forming into DEs and welding to a cylindrical
shell made of the same steel grade. Four DE-cylinder
assemblies have been studied: two 15 mm thick and two
30 mm thick.
Even in the CF state, the impact and fracture toughness
of the welded DE-cylinder assemblies are good. For a CF
DE made from a single plate T27/28J!K50 8C (actual
values K65 to K115 8C). If the CF DE has to be made from
a welded blank T27/28J!K20 8C (actual values K30 to K70 8C) due to the cold forming of the blank welds.
It has been shown that T27/28J can be used to estimate the
fracture toughness parameter T0 in the conventional way.
This, in turn, can be combined with the Master Curve to
calculate the minimum operational temperatures for the
vessel. Because of the toughness reserve of the P420ML2
grade and the modest increase in T27/28J and T0 caused by
forming and ageing, safe minimum operational tempera-
tures are low even with high applied stresses (80% SMYS).
Post-forming heat treatment will not normally be
necessary but it can be used, for example, to improve
the toughness of CF blank welds.
The impact toughness of the CF DE can be determined
from as-delivered plate by simply compressing small plate
samples 15% in plane strain and ageing 30 min at 250 8C
before making Charpy impact specimens. In this way, it is
possible to verify that the TMCP grade to be used meets
toughness requirements in all parts of the DE. The same
approach could also be used for any blank welds.
The upper-shelf toughness of P420ML2 sustained its
considerably high level of 250–300 J, regardless of the
material’s condition, remaining almost unchanged during
cold forming and ageing. On the basis of the high Charpy
upper-shelf values and the results of the notch bending tests,
ductile fracture is not going to be a concern in the
investigated P420ML2 steel in any of its CF and aged
conditions.
The yield and tensile strengths of a CF (or PFHT) DE are
much greater than those of the vessel cylinder at both 20 and
300 8C. Therefore, much higher stresses are required to cause
plastic deformation in the DE than in the cylinder. On the
other hand, the membrane stresses in the cylinder are higher
than in the DE. Therefore the dimensioning of the pressure
vessel is based on the properties of the cylinder (i.e. base
plate) and the safety margin against overload is much greater
in the CF DE than in the cylinder. The elevated Y/T values in
the CF DE close to unity are combined with unreduced
material ductility and are therefore fully acceptable.
Secondary bending stresses are greatest at the DE
knuckle but even CF weld metal, which has the lowest
reduction of area to fracture, has been shown to possess very
good ductility in plane strain bending, which should be more
than sufficient to accommodate the bending strain.
References
[1] Camus M, Desplaces G, Trochain A. Nouveaux reservoirs propane
petit vrac enterres en acier thermomecanique. AFIAP (Fr Assoc
Pressure Vessel Engrs) 1998;October.
[2] Nevasmaa P, Kortelainen O, Wallin K, Porter D. Fracture toughness
assessment and HAZ strain ageing resistance of thermomechanically
processed steel subjected to cold forming before welding Proceedings
of the sixth International Offshore and Polar Engineering Conference,
Los Angeles, 26–31 May 1996, vol. IV. USA: International Society of
Offshore and Polar Engineers (ISOPE); 1996 p. 154–9.
[3] Chapter 3.4 in the background documentation of the ECOPRESS
project available through the site: www.ecopress.org
