ISO/TR 23602:2005
(Main)Toughness of chain steels
Toughness of chain steels
ISO/TR 23602:2005 presents an investigation to quantify the required material toughness for the operational safety of round steel chains.
Résistance des aciers pour chaînes
General Information
Standards Content (Sample)
TECHNICAL ISO/TR
REPORT 23602
First edition
2005-07-01
Toughness of chain steels
Résistance des aciers pour chaînes
Reference number
©
ISO 2005
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ii © ISO 2005 – All rights reserved
Contents Page
Foreword. v
1 Scope .1
2 Normative references .1
3 Types of chain and extraction of specimens.2
4 Materials, chemical composition and heat-treatment.2
5 Tensile tests on chains .2
6 Conventional tensile tests on standard specimens.3
7 Notch impact tests.3
8 Fracture mechanics tests on notched chain links with slits.4
9 Fracture mechanics tests on notched three-point bend specimens with slits and fatigue
cracks.6
10 Correlation between the load-bearing capacity of chain links containing slits and material
toughness.7
11 Fracture mechanics derivation of chain links .7
12 Load-bearing and brittle fracture transition concept — Requirements.8
13 Correlation of test results with data from the literature .9
14 Summary.9
Figure 1 — Results C tests: Standard specimen (EN 10045).11
v
Figure 2 — Load-time diagram for instrumented C .13
v
Figure 3 — Crack arrest load as function of temperature .14
Figure 4 — Noncrystalline area in C — Test.15
v
Figure 5 — Load–COD diagram for TPB specimens with fatigue crack and eroded slits material
type T, T = –40 °C .16
Figure 6 — Load–COD diagram for TPB specimens with fatigue crack and eroded slits material
grade VH, T = –40 °C.17
Figure 7 — Comparison of pre-cracked and eroded chain link material grade VH, T = –40 °C .18
Figure 8 —Chain link specimen .19
Figure 9 — Clip gauge fixed in chain link.20
Figure 10 — Calibration of clip gauge for COD measurement.21
Figure 11 — Test specimen at –40 °C, three links.21
Figure 12 — Effect of slit size on loadability of chain, T = –40 °C .22
Figure 13 — Load–COD diagram for tension tests on chains with eroded slit, T = –40 °C, a/d = 0,22.23
Figure 14 — Load–COD diagram for tension tests on chains with eroded slit, T = –40 °C, a/d = 0,43.24
Figure 15 — Load–COD diagram for tension tests on chains with eroded slit, T = –40 °C, a/d = 0,64.25
Figure 16 — Stable crack growth chain Type T, a/d = 0,22 — Marked by heat tinting.26
Figure 17 — Effect of slit size on absorbed energy of chains, T = –40 °C .27
Figure 18 — Fracture mechanics tests on TPB specimens. 28
Figure 19 — Fracture mechanics concepts. 29
Figure 20 — Fracture toughness values for TPB specimens with eroded slits, T = –40 °C, a/w = 0,5. 30
Figure 21 — R-curves material Type T, T = –40 °C, a/w = 0,4 . 31
Figure 22 — R-curves material Type T, T = –40 °C, a/w = 0,4 . 32
Figure 23 — Nominal fracture stress in chain link, correlated with fracture mechanics properties. 33
Figure 24 — Nominal fracture stress in chain link correlated with fracture mechanics properties. 34
Figure 25 — Correlation of fracture load and C toughness . 35
v
Figure 26 — Correlation nominal fracture stress of chain and C toughness. 36
v
Figure 27 — Finite element analysis of chain links with eroded slit, a/d = 0. 37
Figure 28 — Finite element analysis of chain links with eroded slit, a/d = 0,22. 37
Figure 29 — Finite element analysis of chain links with eroded slit, a/d = 0,64. 38
Figure 30 — Nominal stresses in cracked chain link . 39
[5]
Figure 31 — Correction factor for long bending bars . 40
Figure 32 — Stress intensity factor for long bending bars . 41
Figure 33 — Safety relations between requirement and service conditions . 42
Figure 34 — Nominal fracture stress in chain link, correlated with fracture mechanics properties. 43
Figure 35 — Correlation nominal fracture stress of chain and C toughness. 44
v
Figure 36 — Correlation of fracture and C toughness pre-cracked TPB specimens . 45
v
Table 1 — Chemical composition of the chain steels . 46
Table 2 — Tensile test results on chain specimens (5 links) . 46
Table 3 — Tensile test results on standard specimens (EN 10002, B6*30) T = –40 °C . 46
Table 4 — Results of C tests.47
v
Table 5 — Crack arrest loads of C tests . 47
v
Table 6 — Noncrystalline area of C specimens. 47
v
Table 7 — Transition behaviour of chain materials. 47
Table 8 — Maximum loads of chain-tests at –40 °C . 48
Table 9 — Maximum displacement for chain tests, T = –40 °C. 48
Table 10 — Absorbed energy of the chain tests, T = –40 °C . 48
Table 11 — Evaluation of fracture mechanics test on TPB specimens . 49
Table 12 — Fracture mechanics tests on TPB specimens with eroded slits. 49
Table 13 — Fracture mechanics tests on TPB specimens with fatigue cracks, T = –40 °C . 49
Table 14 — R curves data material Type T . 50
Table 15 — Calculation of K values for chain grade VH . 50
Q
Bibliography . 51
iv © ISO 2005 – All rights reserved
Foreword
ISO (the International Organization for Standardization) is a worldwide federation of national standards bodies
(ISO member bodies). The work of preparing International Standards is normally carried out through ISO
technical committees. Each member body interested in a subject for which a technical committee has been
established has the right to be represented on that committee. International organizations, governmental and
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International Electrotechnical Commission (IEC) on all matters of electrotechnical standardization.
International Standards are drafted in accordance with the rules given in the ISO/IEC Directives, Part 2.
The main task of technical committees is to prepare International Standards. Draft International Standards
adopted by the technical committees are circulated to the member bodies for voting. Publication as an
International Standard requires approval by at least 75 % of the member bodies casting a vote.
In exceptional circumstances, when a technical committee has collected data of a different kind from that
which is normally published as an International Standard (“state of the art”, for example), it may decide by a
simple majority vote of its participating members to publish a Technical Report. A Technical Report is entirely
informative in nature and does not have to be reviewed until the data it provides are considered to be no
longer valid or useful.
Attention is drawn to the possibility that some of the elements of this document may be the subject of patent
rights. ISO shall not be held responsible for identifying any or all such patent rights.
ISO/TR 23602 was prepared by Technical Committee ISO/TC 111, Round steel link chains, chain slings,
components and accessories, Subcommittee SC 1, Chains and chain slings.
TECHNICAL REPORT ISO/TR 23602:2005(E)
Toughness of chain steels
1 Scope
The objective of the investigation presented in this Technical Report was to quantify the required material
toughness for the operational safety of round steel chains. The determination of this toughness must take
place under unfavourable boundary conditions, i.e. at the lowest permissible operating temperature on
damaged (precracked) chains. On the basis of characteristic properties determined on chains and three-point
bend (TPB) specimens and the notch impact energy, a load-bearing and brittle fracture transition temperature
concept was elaborated. From the correlation of the fracture stress of chains, fracture mechanics
characteristic properties and the notch impact energy, minimum values of the notch impact energy can be
derived which yield adequate safety against fracture for the damaged chain. Furthermore, in order to exclude
brittle failure of chains, the position of the lowest permissible operating temperature relative to the NDT
(nil-ductility transition) temperature was determined. Finally, the linkage of the determined fracture mechanics
properties to notch impact energy values and the definition of the requirements for these values furnishes a
simple test and acceptance procedure for chain materials and chain types.
2 Normative references
The following referenced documents are indispensable for the application of this document. For dated
references, only the edition cited applies. For undated references, the latest edition of the referenced
document (including any amendments) applies.
ISO 3077:2001, Short-link chain for lifting purposes — Grade T, (types T, DAT and DT), fine-tolerance hoist
chain
1)
ISO 16872:— , Short link chains for lifting purposes — Grade VH, fine tolerance for manually operated chain
hoists
1)
ISO 16877:— , Short link chains for lifting purposes — Grade TH, fine tolerance for manually operated chain
hoists
DIN 17155, Steel for welded round links; technical delivery conditions
EN 10083-3, Quenched and tempered steels — Part 3: Technical delivery conditions for boron steels
EN 10002 (all parts), Metallic materials — Tensile testing
EN 10045 (all parts), Metallic materials — Charpy impact test
BS 5762, Methods for Crack Opening Displacement (COD) Testing
ASTM E 813, Standard Test for Jic, A Measure of Fracture Toughness
ASTM E 399, Standard Test Method for Plane-Strain Fracture Toughness of Metallic Materials
1) Under preparation.
3 Types of chain and extraction of specimens
For the investigation, chains of size 16 × 48 were selected. The diameter of 16 mm and the length of the
straight shank are suitable for the extraction of notch impact bend and three-point bend specimens for fracture
mechanics tests, as well as small standard tensile test specimens. For fracture mechanics tests of the chains
themselves, the unwelded shank was notched on the inner side. The test cross-section is thus the same for
the chain specimen, the notch impact bend specimen and the standard tensile specimen — which means that
in all cases the testing is based on the same material and heat-treatment conditions.
2)
The 16 × 48 chains were manufactured to conform with ISO 3077:2001, Type T, ISO 16877:— , Grade TH
2)
and ISO 16872:— , Grade VH. Thus chains which may be employed in all hoists (Type T) and exclusively in
hand-operated hoists (Grade TH and Grade VH) are dealt with. Corresponding to their designation, the chains
have a minimum nominal fracture stress of 800 and 1000 N/mm .
4 Materials, chemical composition and heat-treatment
The materials requirements for the manufacture of the chains involved in the investigation are given in Table 1.
Chains conforming to ISO 3077, Type T shall have the minimum contents given in Table 1 as regards the
alloying elements Ni and Cr and/or Mo. Furthermore, the Al and the upper limits for P and S content are
specified.
For both chain types for use in hand operated hoists merely an Al content > 0,025 and limits of P and S to
0,025 and 0,020 respectively are called for. In the lower part of Table 1, the product analyses for the individual
chain specimens are given. A comparison with the specified values shows that all materials meet the
requirements of the corresponding International Standard in every respect. For the material investigated for
chains conforming to ISO 3077, a triple alloyed steel designated 23MnNiCrMo53 in DIN 17115 is invoked. The
two chains of Types TH and VH were made of a manganese-boron steel of European manufacture. The steel
is 19MnB4+Cr conforming to EN 10083-3.
All of the chains investigated underwent an induction heat-treatment. They were austenitised inductively at
1 050 °C and water quenched. Subsequently, conventional tempering took place in an air circulation furnace
under the following conditions:
ISO 3077:2001, Type T: 440 °C/ 2 h
2)
ISO 16877:— , Grade TH: 365 °C/ 2 h
2)
ISO 16872:— , Grade VH: 180 °C/ 2 h
5 Tensile tests on chains
The round steel chains selected for the tests had not been subjected to the manufacturing proof force, by
which the chains are stressed beyond their elastic limit such that in certain regions yield occurs and residual
stresses arise. In subsequent tests, these residual stresses would have affected the determined material
characteristic properties and especially their dependence on notch depth, in a manner which would not have
been able to be understood. For this reason, the application of the manufacturing proof force was dispensed
with. The tensile tests were carried out on the one hand at room temperature with 5-link specimens as in the
corresponding chain standards and on the other hand at the lowest permissible operating temperature of
–40 °C. Since these low-temperature tests were carried out with the same gripping arrangements as with the
fracture mechanics tests on notched and slit specimens, the testing was performed at –40 °C on 3-link
specimens. This is of no importance, since the total ultimate elongation at fracture values is irrelevant due to
the missing application of the manufacturing proof force. They were carried out nevertheless in order to obtain
a starting point as to what deformation capacities the individual chains possessed.
2) Under preparation.
2 © ISO 2005 – All rights reserved
Table 2 gives the requirements for breaking force at room temperature in accordance with the appropriate
standards and the fracture stress resulting therefrom as well as the requirements for total ultimate elongation
A. From this it can be seen that the requirements for fracture stress of chain to ISO 3077, Type T and chain to
2) 2 1)
ISO 16877:— , Grade TH, are greater than 800 N/mm and for chain to ISO 16872:— , Grade VH, greater
than 1 000 N/mm .
The required minimum total ultimate elongation values range from > 10 % for the Type T chain to > 17 % for
the Grade VH chain. In the lower part of Table 2 the values determined on the chains are given. All the
minimum requirements are exceeded by a large margin. At room temperature the Type T chain attained a
2 2
fracture stress of 902 N/mm , while the Grade TH and Grade VH chains reached 806 N/mm and
1 034 N/mm , respectively. This indicates that the strength of the Grade TH and VH chains made of
manganese-boron steel only slightly exceeds the required values. The total ultimate elongations of between
33 % for the Type T chain and 40 % for the Grade TH chain show that the deformation capacity of the
individual chain links is relatively large. The fracture stress values determined at –40 °C do not differ
significantly from those at room temperature for the Type T and Grade TH chains. The Grade VH chain shows
a fall-off which is, however, entirely attributable to the surface. On removal of the –40 °C cold chain from the
cold chamber, the surface of the chain immediately became coated with hoarfrost. This frost led to significant
reduction of the friction between the chain links. This effect influences the fracture stress of the chain as the
material strength increases.
The core hardness values of the chains tested are also given in Table 2. In all cases these correspond with
the determined chain strengths. In all the tensile tests at both test temperatures, the chains failed in the
rounded end.
6 Conventional tensile tests on standard specimens
For the performance of tensile tests conforming to EN 10002, tensile specimens of size 6 × 30 were taken
from the unwelded shanks of the chain links. The specimens were short-proportioned with threaded heads.
These specimens were tested at the lowest permissible operating temperature of –40 °C. Almost identical
results resulted for the 0,2 % yield stress R and ultimate tensile strength R for both Type T and Grade TH
p0,2 m
2 2
chains. The yield stress is about 1 150 N/mm and the tensile strength R about 1 240 N/mm , see Table 3.
m
The same applies for the deformation values, the elongation % A and the reduction of area Z where the A
5 5
values are about 14 % and the Z values at about 65 %. As may be expected, the material of the Grade VH
chain deviates from the above showing a higher 0,2 % yield stress of 1 260 MPa and a tensile strength R of
m
1 430 MPa. Corresponding to the higher strength here, the elongation at fracture is 12 % and the reduction of
area 60 %.
The ratios of the material tensile strength to the breaking strength of the chain lie at mean values of 1,46 for
Grade TH and Type T chains and 1,36 for Grade VH chain. Accordingly, this correlation varies within the
[1]
usual bounds .
7 Notch impact tests
For the performance of the notch impact tests ISO V (Charpy-V), specimens conforming to EN 10045 or
[7]
ISO/R 442 with V notches were taken from the unwelded shanks of the chain links. The tests were carried
out over a temperature range of –80 °C to +80 °C. During every impact test a record of the force-time history
was obtained by means of an instrumented 300 J pendulum impact testing machine to the striker of which
resistance strain gauges were attached. The results are given in Table 4 and are shown in Figure 1 as K -
V
temperature curves. All 3 chain types tested showed a marked transition temperature behaviour. For the
specimens from Type T chain, the transition temperature range lies between –40 and room temperature. For
both the manganese-boron steel chains, however, the transition range is shifted to higher temperatures of
about 0 °C to +80 °C.
On the upper shelf, specimens of Type T chain reached values of about 105 J, those of Grade TH about 80 J,
and those of Grade VH about 55 J. The determined values show good agreement with those obtained for the
same and similar materials and in published results [1] and [2].
More precise information on the brittle fracture transition behaviour of the individual materials can be obtained
from the instrumented notch impact test. Correlations of the crack arrest force in the notch impact test with the
NDT temperature determined in the Pellini drop-weight test show that the temperature at which a crack arrest
[3],[4]
force of 4 kN occurs corresponds to a wide extent to the NDT temperature . Examples of force–time
diagrams are shown in Figure 2. According to Figure 2a), no crack arrest occurs at a test temperature of
–40 °C in a Type T specimen. On raising the test temperature by 10 K, crack arrests occur at a force of 4 kN,
see Figure 2b). If, on the other hand, the specimen is tested on the upper shelf, no rapid crack growth occurs.
This may be seen from Figure 2c) at a test temperature of 20 °C.
The evaluation of all the notch impact tests carried out in the transition region for the 3 materials investigated
are summarised in Table 5 and Figure 3. From the curves NDT temperatures of –30 °C, +17 °C and +22 °C
result respectively for the Type T, Grade TH and Grade VH materials. Consequently, the crack arrest force
and thus the NDT temperatures of the Grade TH and VH chains do not differ significantly. They lie at room
temperature, which means that cracked chains of these qualities can only just arrest cracks.
Another method of determining the brittle fracture transition temperature can also be carried out without
instrumentation of the impact testing machine. In this, an evaluation is made of the fracture surfaces of notch
impact specimens with regard to the proportion of ductile fracture. This proportion is related as a percentage
of the total fracture surface and plotted against test temperature. The criterion for the brittle fracture transition
is 50 % ductile fracture (FAT temperature).
The results for the investigated chains assessed in this manner are +30 °C for Grades TH and VH and –25 °C
for Type T, see Table 6 and Figure 4.
In Table 7 the determined brittle fracture transition temperatures for all the chains using both methods are
summarised. The values for the Type T chain show relatively good agreement of the NDT and FAT
temperatures, whereas the NDT temperatures for the Grade TH and VH chains lie about 10 K lower than the
FAT temperatures. The more reliable method for determination of the brittle fracture transition temperature is
the instrumented notch impact test. Here the NDT temperature is determined by means of objective
characteristic values. The FAT temperature, on the other hand, rests on the subjective estimation of the brittle
or ductile fractions of the fracture surface.
The NDT temperatures of +17 and +22 °C determined for the manganese-boron steels agree exactly with the
NDT temperature determined in earlier work on a manganese-boron steel chain of Category 100. Here
[1]
likewise a value of +20 °C was determined from the instrumented notch impact test .
8 Fracture mechanics tests on notched chain links with slits
For the determination of fracture mechanics characteristic properties, 3-link specimens were selected in which
the middle link was notched on the inner side of the unwelded shank and subsequently fatigued. From this
fatigue, cracks with irregular crack fronts resulted. In contrast to the elliptical crack fronts arising in the normal
case, here the crack advanced much further in the centre of the specimen. This crack advance at the
specimen centre was so great in comparison to that at the edge of the specimen that a fracture mechanics
evaluation of the specimens was not possible. Furthermore, detection of the fatigue crack depth was
extraordinarily difficult. In particular, the cut-off of the resonance test equipment via the drop-off in frequency is
ruled out since the chain link forms a closed ring and thus with advancing crack the frequency drop-off is
extremely small. A possible solution for the introduction of fatigue cracks which are amenable to fracture
mechanics evaluation would perhaps be offered by a circumferential notch in the test cross-section with the
use of the potential probe measurement technique for determination of the crack depth.
For the creation of absolutely reproducible relationships with respect to crack depth and crack front, very
narrow slits were introduced into the chain links by wire spark erosion. For this a fine wire with a diameter of
0,15 mm was employed which produced slit widths of 0,25 mm and slit tip radii of 0,125 mm. In metallographic
sections through the slit tip, no thermally induced structural changes could be detected. Because of this slit tip
radius of 0,125 mm, it is to be expected that in tough material conditions, no significant change occurs in
load-bearing behaviour in comparison with a fatigue crack, since under load the crack tips are likewise
considerably blunted (blunting). This expectation was confirmed by three-point bending specimens. With
approximately the same crack depth to width ratios, the same force/COD plots were obtained.
4 © ISO 2005 – All rights reserved
The steeper path and higher peak load of the curve determined on the specimen with the fatigue crack can be
attributed to the smaller crack depth, see Figure 5. Although, by the use of eroded slits in place of fatigue
cracks with brittle material behaviour, higher failure loads are to be expected. This was likewise confirmed on
three-point bend specimens of Grade VH material, see Figure 6. It is seen that in the brittle material condition,
the failure load of the fatigue crack lies 20 % below that of the eroded slit. The differing gradients of the two
load COD curves are again attributable to the differing crack depths. If one corrects for the crack depth, an
approximately 25 % difference in the fracture load results. A comparison of fatigued and slit chain links led to
the same failure force relationships as in the bending specimens. Here the breaking force of the fatigued chain
link also lies 25 % below that of the chain link with a slit, see Figure 7.
This comprehensible agreement or reproducible deviation in all cases justifies the introduction of eroded slits
in the test chains and permits a later correction of the results. For the systematic investigation of the influence
of defect depth, 3 different defect depths were chosen. The depths of the eroded slits were 3,5 mm, 7,0 mm
and 10,5 mm. These correspond to crack depth-diameter ratios of 0,22, 0,43 and 0,64, see Figure 8. For the
determination of the crack opening on the inner side of the link shank, a COD clip gauge with applied
resistance strain gauges was fastened to the chain link, see Figure 9. The relationship between the clip gauge
opening and strain was calibrated using a micrometer screw, see Figure 10. Figure 11 shows a three-link
chain specimen after the test at –40 °C, which shows the hoarfrost coating mentioned above. In conducting
the tests, the force, testing machine displacement and crack opening were recorded. Examples of force–
displacement and force–crack opening diagrams are shown in Annex 1.
The maximum loads as a function of the particular slit depth and the related elongations determined in the
tests are summarised in Tables 8 and 9 respectively. The listed breaking force values are the averages of
4 individual values. Furthermore, the standard deviations for the particular 4 values are given. As is to be
expected, the material with the highest notch impact energy, Type T, shows the smallest scatter, i.e the
smallest standard deviation. It is noteworthy that the standard deviations of the breaking force values for all
the materials are at a minimum for the crack depth ratio of 0,42. The overall relatively small standard deviation
confirms the reproducibility of the eroded slits with respect to depth and shape.
The relationship between maximum load and relative defect depth is shown in Figure 12. Starting from the
values of the defect-free chain, the maximum load falls off very sharply with a relative slit depth of 0,22, then
with greater relative defects assumes an almost asymptotic path to specific value. This hyperbolic curve form
can be attributed to the supporting effect of the second undefected shank of the chain. It is striking that for the
chain specimens from the manganese-boron steel with the relative slit depth of 0,22 the maximum load has
already fallen to half that for the slit-free specimen.
At larger relative slit depths the maximum loads of the chains of manganese-boron steel are only half those of
Type T chain. Furthermore, it should be stressed that the curves of Grades TH and VH differ only to an
insignificant extent. The small differences in the maximum load are even more astonishing if account is taken
of the great difference in the material strength of the two chains and the widely different tempering
temperatures. Both chains were hardened in the martensitic state. By contrast, the tempering temperatures of
180 °C for Grade VH and 365 °C for Grade TH differed widely. On the other hand, the notch impact energy of
both material conditions is comparable. Thus the toughness is the dominant defining factor for the attainable
maximum load of cracked or notched chains.
Typical Load–COD diagrams for the 3 slit depths and the 3 chains are shown in Figures 13, 14 and 15. These
curves show that for Type T chain failure occurred by overload (plastic collapse). In order to check whether
the collapse failure was preceded by crack initiation and stable crack growth, several chains were loaded up
to differing maximal loads and the resulting stable crack growth therefore was detected by heat tinting (see
example in Figure 16). For this reason, the fracture mechanics concept based on the COD and the J-integral
were used for the quantification of the toughness behaviour of the Type T chain (see Clause 9). In contrast to
this collapse behaviour, the fracture in the Grade TH and VH chains occurred in the linear elastic region.
Consequently, the failure of these chains could be described by linear elastic fracture mechanics.
Along with the force and extension, the tensile tests on the chains were also evaluated with respect to the
absorbed energy at fracture. These results are given in Table 10 and represented graphically in Figure 17.
Here a hyperbolic curve of energy against relative slit depth also results in a drastic fall in energy which may
be seen between the undamaged condition and the 0,22 relative slit depth. The energy is reduced by a factor
of 20 for the Grade TH chain, by a factor of 25 for the Grade VH chain and by a factor of 4 for the Type T
chain. Thus the absorbed energy at fracture at the relative slit depth of 0,22 in the Type T chain is three times
that in the Grade VH and TH chains. This relationship of a factor of 3 between the manganese-boron steel
chains and the Type T chain also remains at greater relative slit depths. It is also noteworthy that at a relative
slit depth of 1 the energy does not reduce to zero because the second shank takes over the load.
9 Fracture mechanics tests on notched three-point bend specimens with slits and
fatigue cracks
9.1 Testing and evaluation technique
In the specimens having the dimensions of notch impact specimens as described in Clause 3, both spark-
eroded slits and fatigue cracks with a crack depth ratio a/w = 0,5 were introduced. During the tests, the
functions, force–bending deflection and force-crack opening (COD) were recorded. The specimen dimensions
and testing arrangements can be seen in Figure 18. In addition, the geometry of the eroded slits is shown.
The test evaluation took place in accordance with the COD and the J integral concepts, see Figure 19. Here
the relationships between force and COD conforming with BS 5762 and between force and bending deflection
in accordance with ASTM E 813 are shown schematically. In addition, the values V and U required for the
p pl
plastic evaluation are plotted. The formulae for the calculation of the fracture mechanics characteristic
properties are given in Table 11. The elastic components δ and J are calculated with the aid of stress
pl el
intensity factor K. The plastic components result from the plastic crack opening – (V ) and energy
I p
components (U ) which can be determined in correspondence with Figure 19. The total values δ and J
pl tot tot
result from the addition of the elastic and plastic components.
9.2 Tests on three-point bend (TPB) specimens with eroded slits
A three-point bend specimen was taken from each of the three types of chain investigated and tested at
–40 °C. Both specimens from Grade TH and Grade VH chains failed in the linear elastic region while the
specimen from Type T chain showed plastic collapse, see Figures 5 and 6. The fracture mechanics
characteristic values calculated in accordance with the relationships in Table 11 are given in Table 12.
K
Q
Ba,,(W −a)W2,5
R
p0,2
In ASTM E 399 two validity criteria are required to be met for linear elastic fracture mechanics tests. The
specimens from Grade TH and VH chains meet the 5 % tangent criterion. However the size condition is barely
3/2
fulfilled. For both types of chain, a K value of about 3 000 N/mm can be calculated. Corresponding to the
Q
largely linear elastic failure the J and δ values are very small. It emerges that the values of the specimen
tot tot
from the Grade VH chain with 50 compared to 45 N/mm for J and 0,033 to 0,024 mm for δ lie slightly
tot tot
higher. In the specimen from the Type T chain, the evaluation relates to the highest load point which more or
less equates to stable crack initiation. The evaluation leads to a J value of 149 N/mm and a δ value of
tot tot
0,106 mm.
The relationship between the crack tip opening (δ ) and the J integral (J ) is shown in Figure 20. There is an
tot tot
almost linear relationship between the two criteria. The sequence with respect to their toughness behaviour, of
the chains investigated, becomes clear from the 3 to 4 times higher values of J and δ for Type T chain.
For the assessment of the effect of the eroded slits on the failure of the specimens, supplementary fatigued
three-point bend specimens were tested at –40 °C. An overview of the properties so determined is given in
Table 13. For the specimens from Grade TH and VH chains the K values obtained from the eroded
Q
3/2 3/2
specimens are reduced from about 3 000 N/mm to about 2 500 N/mm . It is still a matter of the K value
Q
in which here, however, the size conditions are almost fulfilled. As expected, the δ values of the fatigued
specimens are reduced by a factor of about 4 and the J values by a factor of about 2. Due to crack blunting
the J and δ values for the specimen from the Type T chain are again of the same magnitude as for the slit
specimens, see Figure 5.
In the described evaluations of the slit and fatigued specimens from the Type T chain, the peak load was
always used which lies above the crack initiation force and in this case leads to excessive values. In order to
obtain crack initiation, and thus conservative values, the initiation values for J and δ were therefore determined
6 © ISO 2005 – All rights reserved
on fatigued specimens of Type T chain using the multi-specimen technique. In this the extent of the stable
crack growth was determined by use of the heat-tinting technique (see example in Figure 16). A total of three
specimens were used, which were loaded to different levels and then unloaded again. A summary of all the
characteristic properties determined in this manner is given in Table 14. The force–bending deflection and
force–crack opening diagrams recorded in these tests are contained in Annex 2. Specimen 3-7 was loaded
beyond collapse and despite this showed only the relatively small crack growth (∆a) of 0,25 mm. Accordingly,
crack initiation and collapse are almost identical for this material. It should be pointed out that for the
calculation of the U value the specimen indentation (denting) was corrected. Figures 21 and 22 show the
pl
relationship between on the one hand the crack opening (∆a) and, on the other hand, the J integral and crack
tip opening (R curve). The equations for the particular blunting lines result from these standards: in the case of
J, according to ASTM 1820, as J = 2 × σ × ∆a and in the case of CTOD, according to BS 5762 as
Fl
CTOD = 1.4 × ∆a. The parallels to the blunting line through the crack extension ∆a = 0,15 mm give the J and δ
crack initiation values defined in the particular standard. For these values of J = 175 N/mm (Figure 21) and
δ = 0,105 mm (Figure 22) the validity criteria for the appropriate standard are fulfilled. The further
considerations are not based on the crack initiation values as defined above but rather on the real crack
initiation values taken as the intersection of the blunting line and R curve in order to achieve, on the one hand,
absolute conservatism for the tough material and, on the other, compatibility with data from the literature. The
values which result from this as ductile fracture characteristic properties are J = 110 N/mm and δ = 0,065 mm,
i i
see Figures 21 and 22.
10 Correlation between the load-bearing capacity of chain links containing slits and
material toughness
The relationship between the J integral and CTOD, and the gross area nominal stress of the chain containing
slits, is shown in Figures 23 and 24. The gross area nominal stress results from the load at fracture divided by
the sum of the gross cross sectional area of both of the chain link shanks.
Regressively derived families of curves linked the test points. This pattern results, at least in the elastic region,
from the following expressions:
K σ ⋅Π ⋅ay⋅
I
J==
E E
From this follows, according to the stress solved for constant crack depth, a:
J
σ =
c
where c is a constant.
The plotted fracture mechanics characteristic values are those which were determined on the three-point bend
specimens with fatigue cracks. The J integral and CTOD values for the Type T chain are the initiation values
taken from the R curves. With these diagrams in Figures 23 and 24 a link is made between the fracture
mechanics characteristic properties determined on standard specimens and the load-bearing capacity of the
damaged chain in the three quality classes investigated. From the above, the necessary fracture mechanics
properties as a function of the particular defect depth for a required fracture stress of the chain can be
successfully derived.
Corresponding with the objective of the investigation a similar relationship results in the correlation of the
breaking load of the damaged chains with the notch impact energy (see Figure 25). The breaking load can be
converted into the gross area nominal fracture stress from the relationship noted earlier (see Figure 26). Also
from these graphs the necessary notch impact energy for a required load bearing capacity of the damaged
chain with a given defect depth can be derived.
11 Fracture mechanics derivation of chain links
The computation of the stress intensity factor of a cracked component fracture mechanics requires the
relationship between the defect size and the gross area nominal stress. In the case of a round steel chain
which is damaged on one side, i.e. on one shank, the derivation of the gross area nominal stress is not
possible without something more, since in this case it concerns a statically indeterminate system in which it is
not known how the applied force is distributed between the two shanks. In particular, the changing tensile and
bending component of the stress with increasing crack depth is also unknown. To resolve these problems a
systematic analysis using the finite element method (FEM) was performed using the ANSYS program. From
this FE
...
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