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Failure analysis of large press die holder
Failure analysis of large press die holder Hongxun Wang, Peng Jiang, Weifang Zhang, Yaozhong Zhang, Tong Song PII: DOI: Reference:
S1350-6307(15)30109-6 doi: 10.1016/j.engfailanal.2015.10.009 EFA 2711
To appear in: Received date: Revised date: Accepted date:
6 April 2015 7 October 2015 18 October 2015
Please cite this article as: Wang Hongxun, Jiang Peng, Zhang Weifang, Zhang Yaozhong, Song Tong, Failure analysis of large press die holder, (2015), doi: 10.1016/j.engfailanal.2015.10.009
This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
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Failure analysis of large press die holder , Yaozhong
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Zhang c, Tong Song b
Science & Technology on Reliability & Environment Engineering Laboratory,
Beihang University, Beijing, 100191, China
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a
a,
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Hongxun Wang a, Peng Jiang b, Weifang Zhang
Beijing Research Institute of Mechanical & Technology, Beijing,100083,China
c
Wuxi Turbine Blade Co., Ltd., Wuxi, 214174, China
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b
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Abstract
This paper studied the cracking mode and mechanism of the large press die
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holder. Detailed investigations including macroscopic examination, metallographic
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observation, microfractography, chemical analysis and mechanical properties analysis were carried out. The investigations reveal that the failure mechanism of the die
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holder is fast brittle fracture caused by high stress under the condition of material embrittlement. The main crack originated from the fillet of the die holder. The serious
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impact toughness degeneration at the bottom of the die holder resulted in the material embrittlement. The transmission electron microscopy (TEM) investigation indicates that the material properties degeneration is related to the coarsening precipitates of M23C6 and M6C. The finite element analysis (FEA) demonstrates that the fillet suffering high stress is the weak part of the die holder. Keywords: Failure analysis; Fatigue; Properties degeneration; Thermal mechanical stress
Corresponding author. Tel: +86 010 82315759;
E-mail address: [email protected] (W.F. Zhang)
ACCEPTED MANUSCRIPT 1. Introduction Press die holder is the core component of large die forging equipment, which
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will bear formidable mechanical impact load and high thermal cycling load in service
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life. In consideration of the extreme working condition and large size, the design and
period is long, and the cost is high
[1, 2]
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manufacture of large die and die holder is quite difficult. Moreover, the production . In order to assure that the die holder work
well during the service life, the hot forging die steel should have outstanding
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comprehensive mechanical properties, such as high strength, hardness, toughness, wear resistance, fatigue resistance and plastic deformation resistance
[3]
. In fact,
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though the material is of excellent properties, the properties degeneration of die and die holder usually happens due to the high thermal cycling during forging procedure .
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[4]
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Even though the failure of hot forging die has been studied for several decades, the reliability of hot forging die is not so satisfactory. The failure of die is still one of
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the most significant factor which leads to the high cost of large forgings. The failure analysis for large hot forging die and die holder is fairly difficult because of the
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following two reasons. Firstly, the influence factors causing hot forging die failure vary and are hard to control. These factors consist of improper die design, die materials, die manufacturing and forging operations
[5]
. Secondly, the failure
mechanisms of hot forging die may be one of thermal fatigue, mechanical fatigue, fracture, wear, corrosion and deflection as well as several kinds coupling of the above mechanisms
[6]
. Therefore, finite element analysis is usually introduced to simulate
the stress distribution in die during failure analysis as well as the design and manufacturing procedure [7]. This paper studied the failure mode and mechanism of the die holder which is typical but draws little attention in recent researches, due to that both the die holder and the fracture surface are very huge and complex.
ACCEPTED MANUSCRIPT 2. Material and experimental procedure The large press die holder is made up of the upper die holder and the lower die
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holder. The upper die holder produced a crack in the forging procedure. As shown in
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Fig. 1 and Fig. 2. The crack extended as long as 2316.7mm in the inner cavity bottom
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and penetrated the bottom and right side. The die holder was made of 55NiCrMoV7 hot forging die steel by forging process. The die holder was austenitized at 850℃ for 15 hours, then quenched in the water for 30 minutes, after that cooled in the oil to
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room temperature. Finally, the die holder was tempered at 620℃ and 560℃ in order.
Fig. 1 The die holder and macroscopic crack appearance.
In order to study the fracture features, different specimens numbered section 1 to
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section 9 were cut off the fracture surface. The different sections are shown in Fig. 2. Detailed analysis including macroscopic examination, metallographic microscopic and scanning electron microscopy (SEM), energy dispersive spectrometer (EDS), chemical analysis and hardness measurement, was performed on the section 1 to section 9. Mechanical properties testing and TEM analysis were carried out on the section 4 and section 6. FEA was conducted to simulate the thermal mechanical stress distribution in the whole die holder.
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3. Experiment results
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3.1. Macroscopic inspection
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Fig. 2 The whole fracture appearance and specimens number.
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Fig. 2 shows the whole fracture surface. The thick radial ridges, which indicate the fast fracture, can be seen obviously on the fracture surface. The radial ridges are
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convergent to the bottom surface of the die holder inner cavity. The fracture surface shows yellow-blue-black feature, which implies that it has suffered high temperature
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and oxidation. Specimens cut from the left side of section 5 were kept at the temperature of 200℃, 250℃, 300℃, 350℃, 400℃ and 450℃ for 40 minutes, respectively. The results indicate that the blue and black areas experienced the temperature of about 350℃ and 450℃. At the fracture surface, a flat and white semicircle arc area with a diameter of 40 mm at section 1 can be observed, which indicates that fatigue fracture occurred in this area. Some radial ridges converge at the fatigue area, as shown in Fig. 3. The fatigue area is 0.02% of the total fracture surface. Detailed observation reveals that there is a surface cracking zone with a thickness of 5 mm to 10 mm on the both sides of the fatigue area as well as section 2, section 4 and section 9, as shown in Fig. 4. Furthermore, the radial ridges in the surface cracking zone and that initiated from the fatigue tend to different direction by an angle of 30 degrees. This phenomenon
ACCEPTED MANUSCRIPT demonstrates that the crack initiated and propagated to different directions. Most of the radial ridges, except for some initiating from the fatigue area, initiated from
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surface cracking zone. There is no surface cracking zone can be seen in section 3.
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Fig. 3 The white semicircle fatigue appearance of section 1.
ACCEPTED MANUSCRIPT Fig. 4 Surface cracking zone: (a) side of section 1; (b) section 2; (c) section 4; (d) section 9.
Section 5 and section 8 show the fracture features of crack propagation anaphase.
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Similar to section 1, the obvious cambered striations demonstrate that fatigue fracture
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occurred in section 5 and section 8, as shown in Fig. 5. However, the fatigue striations
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is not so dense as that seen in section 1. Section 5 and section 8 were severely oxidized because they suffered the high temperature of around 350℃ and 450℃,
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respectively.
Fig. 5 The fatigue striations in the propagation anaphase: (a) section 5; (b) section 8.
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3.2. Microscopic observation and EDS analysis The different specimens from different areas were observed by SEM. Fig. 6 shows the white semicircle fatigue fracture. Judging from the fatigue striations convergence direction, the ellipse with a rough surface in the white semicircle is the fatigue origin. The fatigue origin has a maximum length of 6.55 mm. Observation of the fatigue origin reveals that the fatigue arcs converged to the center of the ellipse, consequently, it is from the ellipse center that crack initiated.
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Fig. 6 The fracture surface features of section 1 converged to the ellipse center.
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In the ellipse center, there is an area with a maximum length of 350 μm and clear boundary which seems like an inclusion. The morphology of the ellipse center is
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shown in Fig. 7 with high magnification. The ellipse was tested by EDS before and
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after that the surface had been rinsed by 10%H2SO4+1%corrosion inhibitors solution. Table 1 shows the energy spectrum results. The energy spectrum analysis turns out
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that there was an oxidation layer on the ellipse surface. After the oxidation layer was rinsed, the energy spectrum results meet the technical requirement and there is no
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defect [8,9].
Fig. 7 The initiation of fatigue origin of section 1.
ACCEPTED MANUSCRIPT Table 1
Elements
C
O
K
Na
Ni
Mn
Cr
V
Si
Before rinse
16.21
43.60
0.70
0.37
—
1.09
0.23
—
—
0.22
After rinse
0.56
3.19
—
—
1.61
0.79
1.07
0.52
0.096
0.26
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Mo
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The energy spectrum results of the ellipse center (wt.%)
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In order to confirm whether the metallographic structure is abnormal, the metallographic specimens through the ellipse center was made (Fig. 8). Fig. 9 is the
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metallographic structure along the vertical direction of the ellipse surface. Fig. 9 reveals that the ellipse center has the same tempered martensite as other areas and has
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a fine grain size. There is no abnormal structure was found. Therefore, it can be
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concluded that the ellipse area is not a defect and the fatigue crack was not caused by
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defect.
Fig. 8 The metallographic specimen of the fatigue initiation area.
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Fig. 9 The metallographic structure along the vertical direction of the ellipse surface: (a) the ellipse center; (b) the back of the ellipse center.
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The specimens of section 2, section 4, and section 9 were cut along the bottom surface of the inner cavity. They have the similar microscopic fracture features. Fig.
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10 shows that the fracture features are convergent to the fillet of the inner bottom
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surface. Therefore, the surface cracking zones are linear origins. The plenty of micro cracks and the brittle quasi-cleavage features in the surface cracking zone indicate that
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the material has high brittleness, as shown in Fig. 11 and Fig. 12.
Fig. 10 The linear origin of surface cracking zone.
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Fig. 11 The micro cracks of surface cracking zone.
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The Charpy V-notch and tensile testing at room temperature were carried out on the material of surface cracking zone, respectively. The results turn out that the impact
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fracture features are the same as surface cracking zone, as shown in Fig. 12. It can be
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concluded that the surface cracking was fast brittle fracture caused by large impact
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load.
Fig. 12 Comparison of the quasi-cleavage features: (a) surface cracking zone; (b) impact fracture.
Section 3 and section 6 are located in the corner of the fracture, both of which are quasi-cleavage fracture, as shown in Fig. 13. But the fracture features have some differences between the two areas and surface cracking zone in features. The cleavage patterns of section 3 and section 6 are more elongate, and the tearing ridges are more
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indicate that the material brittleness or stress decreased when crack propagated here.
Fig. 13 The quasi-cleavage features of section 3 and section 6: (a) section 3; (b) section 6.
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Section 5 and section 8 are the areas where fatigue fracture happened at the end
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of the whole fracture. Both of the fatigue propagation areas have the quasi-cleavage features with clear ductile fracture features which are similar to section 6, as seen in
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Fig. 14. In non-fatigue area near section 5, there is a distinct difference in microscopic features that shown in Fig. 15. It can be seen that the ductile fracture, which resembles
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the tensile fracture, is predominant features, rather than brittle fracture features. The above two reasons reveal that when the crack propagated at this area, the loading rate was not very high.
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Fig. 14 The fracture features of the anaphase fatigues.
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Fig. 15 The ductile fractures are predominant in the non-fatigue area that near section 5.
Fig. 16 displays the micrograph of the specimens at section 7, which was the
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metaphase of crack propagation. The fracture of different areas show basically the
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same quasi-cleavage features. The results, together with the fracture features of other sections, demonstrate that the fast brittle fracture is the main fracture mode of the
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impact load.
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whole fracture. Furthermore, it can be inferred that the crack was formed under high
Fig. 16 The quasi-cleavage in the metaphase of crack propagation
3.3. Material quality and mechanical properties analysis Metallurgical examination was performed at different areas of the die holder,
ACCEPTED MANUSCRIPT respectively. The material chemical composition meets the technical requirement of 55NiCrMoV7. The grain size of the material is in level 7 to 8. The material contains
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dot shaped non-deformable inclusions. Besides, there was no decarbonization found.
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The microstructure is normal with tempered martensite and a small amount of carbides.
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Mechanical properties testing at different temperature were carried out on the specimens of section 4 (section 4 includes section 4-1, 4-2, and 4-3) and section 6,
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respectively. The results are shown in Table 2 and Table 3. The tensile strength (Rm) of both the areas are uniform and meet the technical requirement (Std) at room
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temperature, which is 1000 MPa~1250 MPa. However, the tensile strength decreases with the temperature increasing. When the temperature reaches 150℃ and above, the tensile strength are below 1000 MPa. The Rockwell hardness (HRC) and Charpy
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impact toughness (akv) of section 6 meet the technical requirement. However, both
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the properties of section 4 are obviously below the technical requirement, especially
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the impact toughness, as shown in Table 3. Compared with section 6, the hardness and toughness of section 4 decreased by 11.3% and 74.8%, respectively. It can be concluded that the properties of section 4 degenerated seriously, and the material
Table 2
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brittleness increased seriously.
The tensile test results of section 4 at different temperature Temperature
Rm (MPa)
Rp0.2 (MPa)
A (%)
Z (%)
Room temperature
1039
839
13.5
47
150℃
979
774
14.0
46
250℃
946
735
13.5
54
350℃
931
723
15.0
64
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Std.
2
3
4
5
6
4#
33.0
32.8
33.0
33.0
33.0
33.0
(HRC)
6#
37.7
37.2
37.3
36.6
37.4
akv
4#
7.0
11.3
11.1
8.2
6#
43.6
42.4
34.0
27.8
Hardness
8
9
33.2
33.3
32.7
36.9
37.8
36.9
36.9
7.3
12.8
6.4
11.1
9.4
42.4
31.3
36.3
45.6
32.3
≥25 2
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(J/cm )
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36~40
7
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1
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The Rockwell hardness and impact toughness of section 4 and section 6
3.4. Material properties degeneration analysis
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In order to investigate the mechanism of the material properties degeneration of
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section 4, the specimens cut down from section 4 and section 6 were studied by TEM. The material analysis shows that the microstructure is temper martensite and carbides
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with numerous dislocations. A large amount of claviform and granular cementite distribute in the basic martensite. In the section 4, many hexagonal prism precipitates with the size between 200
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nm and 300 nm were found, as shown in Fig. 17. The precipitates were tested by selected area electron diffraction (SAED) and EDS. The SAED pattern was indexed in the Fig. 17. The pattern is formed by [001] crystal zone axis diffraction. The EDS analysis results of precipitates are shown in Table 4. The above results indicate that the precipitates were the mixture of M23C6 (M: Fe and Cr mainly) and M6C (M: Fe and Mo mainly) with face-centered cubic structure. However, only a few this kind of precipitates were found in the section 6. The coarsening of M23C6 and M6C usually results in the embrittlement of die steel[10-12]. Therefore, it was inferred that the material properties degeneration may be related to the precipitates of M23C6 and M6C. The mechanism how the precipitates affected the mechanical properties need to be further studied.
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Fig. 17 The TEM morphology and indexed diffraction pattern of the precipitates of section 4: (a)
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the morphology of precipitates; (b) the indexed diffraction pattern. Table 4
NiK
wt.%
0.59
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FeK
CrK
MoK
VK
77.46
13.80
7.58
0.57
79.14
15.14
4.51
0.64
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at.%
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Elements
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The energy spectrum analysis results of precipitates
3.5. Finite element analysis of thermal mechanical stress Based on the actual usage, the die holder should be preheated to 200℃ using heating rods before forging. During the forging procedure, the inner bottom of the die holder, which contacts with the die, will reach the thermal equilibrium about 350℃. While the temperature of outer bottom will stay at 200℃. The temperature gradient between the inner bottom and the outer bottom will lead to different thermal stresses at different areas. The matching installation of blade, die, and the die holder is shown in Fig. 18. The turbine blade is forged with a complex surface. The surface with space curve will inevitably produce horizontal lateral force to the die holder. Under the specified service condition, the maximum impact load that the die holder suffered was
ACCEPTED MANUSCRIPT simplified into four equivalent static forces. as shown in Fig. 19. To facilitate the
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calculation, FEA model was simplified as follows [7, 13]:
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Fig. 18 The modeling diagram of the matching installation of blade, die, and the die holder.
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was 200℃.
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(1) The temperature of the inner bottom was set at 350℃, while the outer bottom
(2) The outer bottom of the die holder was fully insulated, while the thermal of other outer surfaces were convective with air.
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(3) As shown in Fig. 19. The inner bottom which contact with the die and the two striking surfaces of the die holder suffered the same normal force of 175 MN.
(4) Considering the torsional force of the blade deformation during forging, the two side walls of the die holder bore the horizontal lateral force of 130 MN.
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Fig. 19 The force diagram of the die holder.
The geometry model of FEA was disassembled by tetrahedron element. The
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element size was 0.01m. At the transitional section and in the hole, the meshes were
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further refined by an element size of 0.001m. Finally, the whole die holder had 4975448 nodes and 3081742 elements.
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The surface stress distribution of the die holder is shown in Fig. 20. Due to the effect of temperature and forging reaction, the stress of the inner bottom surface
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generally reaches 700 MPa. The stress of the center hole even reaches 1400MPa. The stress of the two striking surfaces are up to 600 MPa because of the direct strike load. However, the bottom fillet of the die holder has a high stress of 1010.6 MPa, as shown in Fig. 20. Fig. 21 shows the transverse stresses distribution of the die holder. The stresses of the fillet area are close to 800 MPa.
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Fig. 20 The surface stress distribution of the die holder.
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Fig. 21 The transverse stress distribution of the die holde.
The analysis results show that the fillet of the die holder has a higher stress than that of other areas. Obviously, the stress of the center hole is compressive stress, while that of the fillet is tensile stress. The tensile stress is more likely to lead to the damage of the die holder, relatively. Therefore, the fillet where the fracture happened is the weak part of the die holder.
4. Discussion The failure of the die holder was caused by high stress under the condition of material embrittlement. The white semicircle fatigue was caused by cycling
ACCEPTED MANUSCRIPT compressive stress from the die. The material embrittlement was the result of the serious degeneration of impact
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toughness and hardness. The degeneration of impact toughness and hardness were
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caused by the long-term thermal cycling. Under the regular service condition, the temperature of the die holder could reach 200℃~350℃. However, the die holder was
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preheated by flame whose temperature was difficult to control. The temperature sometimes even reached 500℃ or so. When replacing forging dies, the die holder was
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water-cooled to increase the cooling rate sometimes. Therefore, the thermal cycling appeared and lead to cycling tempering behavior which resulted in the material
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brittlement. Due to the higher thermal cycling, the properties degeneration of the bottom was more serious than other areas. Furthermore, The mechanism of material properties degeneration is inferred to be related to the coarsening precipitates of
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M23C6 and M6C.
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The stress around the fillet where the crack initiated were very high. The highest
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stress (1010.6 MPa) was beyond the material tensile strength (931 MPa) at 350℃. The investigation of the forging procedure reveals that there were many non-standard operations. Firstly, the adjusting wedge, which was used to fix the die, was mounted
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before the die holder being preheated to 200℃. After the temperature reaching 200℃, the thermal expansion would result in large lateral forces to the die holder. Secondly, the strength and hardness of forging dies are lower than required. The low strength and hardness would lead to the deformation of forging dies, and extrude on both sides of the die holder. Therefore, the actual stresses may be higher than 1010.6 MPa. In addition, the die holder with the fatigue crack was equivalent to a notch specimen. Under the high impact load during forging, the crack initiated at the fillet and formed the surface cracking zone. Then the crack rapidly grew. Finally, the crack crossed the bottom of the die holder and formed the brittle quasi-cleavage fracture morphology. When the crack reached section 5 and section 8, with stress and energy decreasing, the crack formed the last fatigue fracture at section 5 and section 8 until it was found.
ACCEPTED MANUSCRIPT 5. Conclusions (1) The failure of the die holder was caused by high stress under the condition of
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material embrittlement.
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(2) The main crack originated from the fillet of the die holder. The surface
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cracking zone is the place where the main crack initiates. The main fracture mode is fast brittle quasi-cleavage fracture.
(3) The serious impact toughness degeneration at the bottom of the die holder
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results in the material embrittlement. The material properties degeneration is related to the coarsening precipitates of M23C6 and M6C.
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(4) The white semicircle fatigue was caused by cycling compressive stresses from the forging die. The fatigue has no connection with the material quality.
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Acknowledgement
This research was supported by national fund (No.JSZL2014601B004). The
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authors wish to thank Wuxi Turbine Blade Co., Ltd. for the technology and equipment support. The guidance and help of Hongxiang Jing is greatly acknowledged.
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References
[1] Bandara CS, Siriwardane SC, Dissanayake UI, Dissanayake R. Fatigue failure predictions for steels in the very high cycle region – A review and recommendations. Eng Fail Anal 2014; 45:421-35. [2] Kchaou M, Elleuch R, Desplanques Y, Boidin X, Degallaix G. Failure mechanisms of H13 die on relation to the forging process – A case study of brass gas valves. Eng Fail Anal 2010; 17: 403-15. [3] Ebara R. Fatigue crack initiation and propagation behavior of forging die steels. Int J Fatigue 2010; 32: 830-40. [4] Maktouf W, Saï K. An investigation of premature fatigue failures of gas turbine blade. Eng Fail Anal 2015; 47:89-101. [5] Ebara R, Kubota K. Failure analysis of hot forging dies for automotive components. Eng Fail Anal 2008; 15: 881-93. [6] D. Papageorgiou C. Medrea, N. Kyriakou. Failure analysis of H13 working die used in plastic injection moulding. Eng Fail Anal 2013; 35: 355-59. [7] Fan MT, Sun MY, Li DZ. Simulation and optimization of heat treatment process for large press die holder. J Mater Eng 2011; 11:50.
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[8] G. Bernhart, G. Moulinier, O. Brucelle, D. Delagnes. High temperature low cycle fatigue behaviour of a martensitic forging tool steel. 1999; 21: 179-86. [9] D. Delagnesa, F. Rézaï-Ariaa, C. Levaillant. Influence of testing and tempering temperatures on fatigue behaviour, life and crack initiation mechanisms in a 5%Cr martensitic steel. 2010; 2: 427-39. [10] N. Mebarki, D. Delagnes, P. Lamesle, F. Delmas, C. Levaillant. Relationship between microstructure and mechanical properties of a 5% Cr tempered martensitic tool steel. Mater Sci Eng 2004; 171-75. [11] Zhou J, Ma DS, Chi HX, Chen ZZ, LI XY. Microstructure and properties of hot working die steel H13MOD. J Iron Steel Res 2013; 20: 117-25. [12] Z. Gronostajski, M. Kaszuba, M. Hawryluk, M. Zwierzchowski. A review of the degradation mechanisms of the hot forging tools. Arch Civ Mech Eng 2014; 14: 528-39. [13] Zhang ZP, Bernhart G, Delagnes D. Cyclic behaviour constitutive modelling of a tempered martensitic steel including ageing effect. Int J Fatigue 2008; 30: 706-16.
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Highlights
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(1) The failure were caused by high stress under the condition of material
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embrittlement.
(2) The fracture morphology and mechanical properties were investigated.
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(3) The mechanism of material properties degeneration was studied. (4) The micro structure and precipitates were studied by TEM.
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(5) The thermal mechanical stress was simulated by finite element analysis.
S1350-6307(15)30109-6 doi: 10.1016/j.engfailanal.2015.10.009 EFA 2711
To appear in: Received date: Revised date: Accepted date:
6 April 2015 7 October 2015 18 October 2015
Please cite this article as: Wang Hongxun, Jiang Peng, Zhang Weifang, Zhang Yaozhong, Song Tong, Failure analysis of large press die holder, (2015), doi: 10.1016/j.engfailanal.2015.10.009
This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
ACCEPTED MANUSCRIPT
Failure analysis of large press die holder , Yaozhong
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Zhang c, Tong Song b
Science & Technology on Reliability & Environment Engineering Laboratory,
Beihang University, Beijing, 100191, China
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a
a,
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Hongxun Wang a, Peng Jiang b, Weifang Zhang
Beijing Research Institute of Mechanical & Technology, Beijing,100083,China
c
Wuxi Turbine Blade Co., Ltd., Wuxi, 214174, China
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b
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Abstract
This paper studied the cracking mode and mechanism of the large press die
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holder. Detailed investigations including macroscopic examination, metallographic
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observation, microfractography, chemical analysis and mechanical properties analysis were carried out. The investigations reveal that the failure mechanism of the die
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holder is fast brittle fracture caused by high stress under the condition of material embrittlement. The main crack originated from the fillet of the die holder. The serious
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impact toughness degeneration at the bottom of the die holder resulted in the material embrittlement. The transmission electron microscopy (TEM) investigation indicates that the material properties degeneration is related to the coarsening precipitates of M23C6 and M6C. The finite element analysis (FEA) demonstrates that the fillet suffering high stress is the weak part of the die holder. Keywords: Failure analysis; Fatigue; Properties degeneration; Thermal mechanical stress
Corresponding author. Tel: +86 010 82315759;
E-mail address: [email protected] (W.F. Zhang)
ACCEPTED MANUSCRIPT 1. Introduction Press die holder is the core component of large die forging equipment, which
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will bear formidable mechanical impact load and high thermal cycling load in service
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life. In consideration of the extreme working condition and large size, the design and
period is long, and the cost is high
[1, 2]
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manufacture of large die and die holder is quite difficult. Moreover, the production . In order to assure that the die holder work
well during the service life, the hot forging die steel should have outstanding
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comprehensive mechanical properties, such as high strength, hardness, toughness, wear resistance, fatigue resistance and plastic deformation resistance
[3]
. In fact,
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though the material is of excellent properties, the properties degeneration of die and die holder usually happens due to the high thermal cycling during forging procedure .
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[4]
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Even though the failure of hot forging die has been studied for several decades, the reliability of hot forging die is not so satisfactory. The failure of die is still one of
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the most significant factor which leads to the high cost of large forgings. The failure analysis for large hot forging die and die holder is fairly difficult because of the
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following two reasons. Firstly, the influence factors causing hot forging die failure vary and are hard to control. These factors consist of improper die design, die materials, die manufacturing and forging operations
[5]
. Secondly, the failure
mechanisms of hot forging die may be one of thermal fatigue, mechanical fatigue, fracture, wear, corrosion and deflection as well as several kinds coupling of the above mechanisms
[6]
. Therefore, finite element analysis is usually introduced to simulate
the stress distribution in die during failure analysis as well as the design and manufacturing procedure [7]. This paper studied the failure mode and mechanism of the die holder which is typical but draws little attention in recent researches, due to that both the die holder and the fracture surface are very huge and complex.
ACCEPTED MANUSCRIPT 2. Material and experimental procedure The large press die holder is made up of the upper die holder and the lower die
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holder. The upper die holder produced a crack in the forging procedure. As shown in
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Fig. 1 and Fig. 2. The crack extended as long as 2316.7mm in the inner cavity bottom
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and penetrated the bottom and right side. The die holder was made of 55NiCrMoV7 hot forging die steel by forging process. The die holder was austenitized at 850℃ for 15 hours, then quenched in the water for 30 minutes, after that cooled in the oil to
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room temperature. Finally, the die holder was tempered at 620℃ and 560℃ in order.
Fig. 1 The die holder and macroscopic crack appearance.
In order to study the fracture features, different specimens numbered section 1 to
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section 9 were cut off the fracture surface. The different sections are shown in Fig. 2. Detailed analysis including macroscopic examination, metallographic microscopic and scanning electron microscopy (SEM), energy dispersive spectrometer (EDS), chemical analysis and hardness measurement, was performed on the section 1 to section 9. Mechanical properties testing and TEM analysis were carried out on the section 4 and section 6. FEA was conducted to simulate the thermal mechanical stress distribution in the whole die holder.
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3. Experiment results
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3.1. Macroscopic inspection
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Fig. 2 The whole fracture appearance and specimens number.
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Fig. 2 shows the whole fracture surface. The thick radial ridges, which indicate the fast fracture, can be seen obviously on the fracture surface. The radial ridges are
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convergent to the bottom surface of the die holder inner cavity. The fracture surface shows yellow-blue-black feature, which implies that it has suffered high temperature
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and oxidation. Specimens cut from the left side of section 5 were kept at the temperature of 200℃, 250℃, 300℃, 350℃, 400℃ and 450℃ for 40 minutes, respectively. The results indicate that the blue and black areas experienced the temperature of about 350℃ and 450℃. At the fracture surface, a flat and white semicircle arc area with a diameter of 40 mm at section 1 can be observed, which indicates that fatigue fracture occurred in this area. Some radial ridges converge at the fatigue area, as shown in Fig. 3. The fatigue area is 0.02% of the total fracture surface. Detailed observation reveals that there is a surface cracking zone with a thickness of 5 mm to 10 mm on the both sides of the fatigue area as well as section 2, section 4 and section 9, as shown in Fig. 4. Furthermore, the radial ridges in the surface cracking zone and that initiated from the fatigue tend to different direction by an angle of 30 degrees. This phenomenon
ACCEPTED MANUSCRIPT demonstrates that the crack initiated and propagated to different directions. Most of the radial ridges, except for some initiating from the fatigue area, initiated from
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surface cracking zone. There is no surface cracking zone can be seen in section 3.
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Fig. 3 The white semicircle fatigue appearance of section 1.
ACCEPTED MANUSCRIPT Fig. 4 Surface cracking zone: (a) side of section 1; (b) section 2; (c) section 4; (d) section 9.
Section 5 and section 8 show the fracture features of crack propagation anaphase.
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Similar to section 1, the obvious cambered striations demonstrate that fatigue fracture
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occurred in section 5 and section 8, as shown in Fig. 5. However, the fatigue striations
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is not so dense as that seen in section 1. Section 5 and section 8 were severely oxidized because they suffered the high temperature of around 350℃ and 450℃,
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respectively.
Fig. 5 The fatigue striations in the propagation anaphase: (a) section 5; (b) section 8.
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3.2. Microscopic observation and EDS analysis The different specimens from different areas were observed by SEM. Fig. 6 shows the white semicircle fatigue fracture. Judging from the fatigue striations convergence direction, the ellipse with a rough surface in the white semicircle is the fatigue origin. The fatigue origin has a maximum length of 6.55 mm. Observation of the fatigue origin reveals that the fatigue arcs converged to the center of the ellipse, consequently, it is from the ellipse center that crack initiated.
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Fig. 6 The fracture surface features of section 1 converged to the ellipse center.
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In the ellipse center, there is an area with a maximum length of 350 μm and clear boundary which seems like an inclusion. The morphology of the ellipse center is
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shown in Fig. 7 with high magnification. The ellipse was tested by EDS before and
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after that the surface had been rinsed by 10%H2SO4+1%corrosion inhibitors solution. Table 1 shows the energy spectrum results. The energy spectrum analysis turns out
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that there was an oxidation layer on the ellipse surface. After the oxidation layer was rinsed, the energy spectrum results meet the technical requirement and there is no
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defect [8,9].
Fig. 7 The initiation of fatigue origin of section 1.
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Elements
C
O
K
Na
Ni
Mn
Cr
V
Si
Before rinse
16.21
43.60
0.70
0.37
—
1.09
0.23
—
—
0.22
After rinse
0.56
3.19
—
—
1.61
0.79
1.07
0.52
0.096
0.26
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Mo
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The energy spectrum results of the ellipse center (wt.%)
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In order to confirm whether the metallographic structure is abnormal, the metallographic specimens through the ellipse center was made (Fig. 8). Fig. 9 is the
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metallographic structure along the vertical direction of the ellipse surface. Fig. 9 reveals that the ellipse center has the same tempered martensite as other areas and has
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a fine grain size. There is no abnormal structure was found. Therefore, it can be
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concluded that the ellipse area is not a defect and the fatigue crack was not caused by
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defect.
Fig. 8 The metallographic specimen of the fatigue initiation area.
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Fig. 9 The metallographic structure along the vertical direction of the ellipse surface: (a) the ellipse center; (b) the back of the ellipse center.
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The specimens of section 2, section 4, and section 9 were cut along the bottom surface of the inner cavity. They have the similar microscopic fracture features. Fig.
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10 shows that the fracture features are convergent to the fillet of the inner bottom
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surface. Therefore, the surface cracking zones are linear origins. The plenty of micro cracks and the brittle quasi-cleavage features in the surface cracking zone indicate that
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the material has high brittleness, as shown in Fig. 11 and Fig. 12.
Fig. 10 The linear origin of surface cracking zone.
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Fig. 11 The micro cracks of surface cracking zone.
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The Charpy V-notch and tensile testing at room temperature were carried out on the material of surface cracking zone, respectively. The results turn out that the impact
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fracture features are the same as surface cracking zone, as shown in Fig. 12. It can be
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concluded that the surface cracking was fast brittle fracture caused by large impact
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load.
Fig. 12 Comparison of the quasi-cleavage features: (a) surface cracking zone; (b) impact fracture.
Section 3 and section 6 are located in the corner of the fracture, both of which are quasi-cleavage fracture, as shown in Fig. 13. But the fracture features have some differences between the two areas and surface cracking zone in features. The cleavage patterns of section 3 and section 6 are more elongate, and the tearing ridges are more
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indicate that the material brittleness or stress decreased when crack propagated here.
Fig. 13 The quasi-cleavage features of section 3 and section 6: (a) section 3; (b) section 6.
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Section 5 and section 8 are the areas where fatigue fracture happened at the end
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of the whole fracture. Both of the fatigue propagation areas have the quasi-cleavage features with clear ductile fracture features which are similar to section 6, as seen in
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Fig. 14. In non-fatigue area near section 5, there is a distinct difference in microscopic features that shown in Fig. 15. It can be seen that the ductile fracture, which resembles
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the tensile fracture, is predominant features, rather than brittle fracture features. The above two reasons reveal that when the crack propagated at this area, the loading rate was not very high.
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Fig. 14 The fracture features of the anaphase fatigues.
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Fig. 15 The ductile fractures are predominant in the non-fatigue area that near section 5.
Fig. 16 displays the micrograph of the specimens at section 7, which was the
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metaphase of crack propagation. The fracture of different areas show basically the
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same quasi-cleavage features. The results, together with the fracture features of other sections, demonstrate that the fast brittle fracture is the main fracture mode of the
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impact load.
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whole fracture. Furthermore, it can be inferred that the crack was formed under high
Fig. 16 The quasi-cleavage in the metaphase of crack propagation
3.3. Material quality and mechanical properties analysis Metallurgical examination was performed at different areas of the die holder,
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dot shaped non-deformable inclusions. Besides, there was no decarbonization found.
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The microstructure is normal with tempered martensite and a small amount of carbides.
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Mechanical properties testing at different temperature were carried out on the specimens of section 4 (section 4 includes section 4-1, 4-2, and 4-3) and section 6,
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respectively. The results are shown in Table 2 and Table 3. The tensile strength (Rm) of both the areas are uniform and meet the technical requirement (Std) at room
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temperature, which is 1000 MPa~1250 MPa. However, the tensile strength decreases with the temperature increasing. When the temperature reaches 150℃ and above, the tensile strength are below 1000 MPa. The Rockwell hardness (HRC) and Charpy
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impact toughness (akv) of section 6 meet the technical requirement. However, both
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the properties of section 4 are obviously below the technical requirement, especially
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the impact toughness, as shown in Table 3. Compared with section 6, the hardness and toughness of section 4 decreased by 11.3% and 74.8%, respectively. It can be concluded that the properties of section 4 degenerated seriously, and the material
Table 2
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brittleness increased seriously.
The tensile test results of section 4 at different temperature Temperature
Rm (MPa)
Rp0.2 (MPa)
A (%)
Z (%)
Room temperature
1039
839
13.5
47
150℃
979
774
14.0
46
250℃
946
735
13.5
54
350℃
931
723
15.0
64
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Std.
2
3
4
5
6
4#
33.0
32.8
33.0
33.0
33.0
33.0
(HRC)
6#
37.7
37.2
37.3
36.6
37.4
akv
4#
7.0
11.3
11.1
8.2
6#
43.6
42.4
34.0
27.8
Hardness
8
9
33.2
33.3
32.7
36.9
37.8
36.9
36.9
7.3
12.8
6.4
11.1
9.4
42.4
31.3
36.3
45.6
32.3
≥25 2
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(J/cm )
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36~40
7
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1
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The Rockwell hardness and impact toughness of section 4 and section 6
3.4. Material properties degeneration analysis
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In order to investigate the mechanism of the material properties degeneration of
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section 4, the specimens cut down from section 4 and section 6 were studied by TEM. The material analysis shows that the microstructure is temper martensite and carbides
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with numerous dislocations. A large amount of claviform and granular cementite distribute in the basic martensite. In the section 4, many hexagonal prism precipitates with the size between 200
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nm and 300 nm were found, as shown in Fig. 17. The precipitates were tested by selected area electron diffraction (SAED) and EDS. The SAED pattern was indexed in the Fig. 17. The pattern is formed by [001] crystal zone axis diffraction. The EDS analysis results of precipitates are shown in Table 4. The above results indicate that the precipitates were the mixture of M23C6 (M: Fe and Cr mainly) and M6C (M: Fe and Mo mainly) with face-centered cubic structure. However, only a few this kind of precipitates were found in the section 6. The coarsening of M23C6 and M6C usually results in the embrittlement of die steel[10-12]. Therefore, it was inferred that the material properties degeneration may be related to the precipitates of M23C6 and M6C. The mechanism how the precipitates affected the mechanical properties need to be further studied.
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Fig. 17 The TEM morphology and indexed diffraction pattern of the precipitates of section 4: (a)
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the morphology of precipitates; (b) the indexed diffraction pattern. Table 4
NiK
wt.%
0.59
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FeK
CrK
MoK
VK
77.46
13.80
7.58
0.57
79.14
15.14
4.51
0.64
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at.%
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The energy spectrum analysis results of precipitates
3.5. Finite element analysis of thermal mechanical stress Based on the actual usage, the die holder should be preheated to 200℃ using heating rods before forging. During the forging procedure, the inner bottom of the die holder, which contacts with the die, will reach the thermal equilibrium about 350℃. While the temperature of outer bottom will stay at 200℃. The temperature gradient between the inner bottom and the outer bottom will lead to different thermal stresses at different areas. The matching installation of blade, die, and the die holder is shown in Fig. 18. The turbine blade is forged with a complex surface. The surface with space curve will inevitably produce horizontal lateral force to the die holder. Under the specified service condition, the maximum impact load that the die holder suffered was
ACCEPTED MANUSCRIPT simplified into four equivalent static forces. as shown in Fig. 19. To facilitate the
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calculation, FEA model was simplified as follows [7, 13]:
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Fig. 18 The modeling diagram of the matching installation of blade, die, and the die holder.
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was 200℃.
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(1) The temperature of the inner bottom was set at 350℃, while the outer bottom
(2) The outer bottom of the die holder was fully insulated, while the thermal of other outer surfaces were convective with air.
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(3) As shown in Fig. 19. The inner bottom which contact with the die and the two striking surfaces of the die holder suffered the same normal force of 175 MN.
(4) Considering the torsional force of the blade deformation during forging, the two side walls of the die holder bore the horizontal lateral force of 130 MN.
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Fig. 19 The force diagram of the die holder.
The geometry model of FEA was disassembled by tetrahedron element. The
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element size was 0.01m. At the transitional section and in the hole, the meshes were
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further refined by an element size of 0.001m. Finally, the whole die holder had 4975448 nodes and 3081742 elements.
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The surface stress distribution of the die holder is shown in Fig. 20. Due to the effect of temperature and forging reaction, the stress of the inner bottom surface
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generally reaches 700 MPa. The stress of the center hole even reaches 1400MPa. The stress of the two striking surfaces are up to 600 MPa because of the direct strike load. However, the bottom fillet of the die holder has a high stress of 1010.6 MPa, as shown in Fig. 20. Fig. 21 shows the transverse stresses distribution of the die holder. The stresses of the fillet area are close to 800 MPa.
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Fig. 20 The surface stress distribution of the die holder.
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Fig. 21 The transverse stress distribution of the die holde.
The analysis results show that the fillet of the die holder has a higher stress than that of other areas. Obviously, the stress of the center hole is compressive stress, while that of the fillet is tensile stress. The tensile stress is more likely to lead to the damage of the die holder, relatively. Therefore, the fillet where the fracture happened is the weak part of the die holder.
4. Discussion The failure of the die holder was caused by high stress under the condition of material embrittlement. The white semicircle fatigue was caused by cycling
ACCEPTED MANUSCRIPT compressive stress from the die. The material embrittlement was the result of the serious degeneration of impact
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toughness and hardness. The degeneration of impact toughness and hardness were
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caused by the long-term thermal cycling. Under the regular service condition, the temperature of the die holder could reach 200℃~350℃. However, the die holder was
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preheated by flame whose temperature was difficult to control. The temperature sometimes even reached 500℃ or so. When replacing forging dies, the die holder was
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water-cooled to increase the cooling rate sometimes. Therefore, the thermal cycling appeared and lead to cycling tempering behavior which resulted in the material
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brittlement. Due to the higher thermal cycling, the properties degeneration of the bottom was more serious than other areas. Furthermore, The mechanism of material properties degeneration is inferred to be related to the coarsening precipitates of
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M23C6 and M6C.
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The stress around the fillet where the crack initiated were very high. The highest
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stress (1010.6 MPa) was beyond the material tensile strength (931 MPa) at 350℃. The investigation of the forging procedure reveals that there were many non-standard operations. Firstly, the adjusting wedge, which was used to fix the die, was mounted
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before the die holder being preheated to 200℃. After the temperature reaching 200℃, the thermal expansion would result in large lateral forces to the die holder. Secondly, the strength and hardness of forging dies are lower than required. The low strength and hardness would lead to the deformation of forging dies, and extrude on both sides of the die holder. Therefore, the actual stresses may be higher than 1010.6 MPa. In addition, the die holder with the fatigue crack was equivalent to a notch specimen. Under the high impact load during forging, the crack initiated at the fillet and formed the surface cracking zone. Then the crack rapidly grew. Finally, the crack crossed the bottom of the die holder and formed the brittle quasi-cleavage fracture morphology. When the crack reached section 5 and section 8, with stress and energy decreasing, the crack formed the last fatigue fracture at section 5 and section 8 until it was found.
ACCEPTED MANUSCRIPT 5. Conclusions (1) The failure of the die holder was caused by high stress under the condition of
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material embrittlement.
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(2) The main crack originated from the fillet of the die holder. The surface
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cracking zone is the place where the main crack initiates. The main fracture mode is fast brittle quasi-cleavage fracture.
(3) The serious impact toughness degeneration at the bottom of the die holder
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results in the material embrittlement. The material properties degeneration is related to the coarsening precipitates of M23C6 and M6C.
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(4) The white semicircle fatigue was caused by cycling compressive stresses from the forging die. The fatigue has no connection with the material quality.
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Acknowledgement
This research was supported by national fund (No.JSZL2014601B004). The
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authors wish to thank Wuxi Turbine Blade Co., Ltd. for the technology and equipment support. The guidance and help of Hongxiang Jing is greatly acknowledged.
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References
[1] Bandara CS, Siriwardane SC, Dissanayake UI, Dissanayake R. Fatigue failure predictions for steels in the very high cycle region – A review and recommendations. Eng Fail Anal 2014; 45:421-35. [2] Kchaou M, Elleuch R, Desplanques Y, Boidin X, Degallaix G. Failure mechanisms of H13 die on relation to the forging process – A case study of brass gas valves. Eng Fail Anal 2010; 17: 403-15. [3] Ebara R. Fatigue crack initiation and propagation behavior of forging die steels. Int J Fatigue 2010; 32: 830-40. [4] Maktouf W, Saï K. An investigation of premature fatigue failures of gas turbine blade. Eng Fail Anal 2015; 47:89-101. [5] Ebara R, Kubota K. Failure analysis of hot forging dies for automotive components. Eng Fail Anal 2008; 15: 881-93. [6] D. Papageorgiou C. Medrea, N. Kyriakou. Failure analysis of H13 working die used in plastic injection moulding. Eng Fail Anal 2013; 35: 355-59. [7] Fan MT, Sun MY, Li DZ. Simulation and optimization of heat treatment process for large press die holder. J Mater Eng 2011; 11:50.
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[8] G. Bernhart, G. Moulinier, O. Brucelle, D. Delagnes. High temperature low cycle fatigue behaviour of a martensitic forging tool steel. 1999; 21: 179-86. [9] D. Delagnesa, F. Rézaï-Ariaa, C. Levaillant. Influence of testing and tempering temperatures on fatigue behaviour, life and crack initiation mechanisms in a 5%Cr martensitic steel. 2010; 2: 427-39. [10] N. Mebarki, D. Delagnes, P. Lamesle, F. Delmas, C. Levaillant. Relationship between microstructure and mechanical properties of a 5% Cr tempered martensitic tool steel. Mater Sci Eng 2004; 171-75. [11] Zhou J, Ma DS, Chi HX, Chen ZZ, LI XY. Microstructure and properties of hot working die steel H13MOD. J Iron Steel Res 2013; 20: 117-25. [12] Z. Gronostajski, M. Kaszuba, M. Hawryluk, M. Zwierzchowski. A review of the degradation mechanisms of the hot forging tools. Arch Civ Mech Eng 2014; 14: 528-39. [13] Zhang ZP, Bernhart G, Delagnes D. Cyclic behaviour constitutive modelling of a tempered martensitic steel including ageing effect. Int J Fatigue 2008; 30: 706-16.
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Highlights
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(1) The failure were caused by high stress under the condition of material
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embrittlement.
(2) The fracture morphology and mechanical properties were investigated.
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(3) The mechanism of material properties degeneration was studied. (4) The micro structure and precipitates were studied by TEM.
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(5) The thermal mechanical stress was simulated by finite element analysis.