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Failure analysis of a large ball valve for pipe-lines
Engineering Failure Analysis 32 (2013) 167–177
Contents lists available at SciVerse ScienceDirect
Engineering Failure Analysis journal homepage: www.elsevier.com/locate/engfailanal
Failure analysis of a large ball valve for pipe-lines S. Bagherifard a, I. Fernández Pariente b,⇑, M. Guagliano a a b
Politecnico di Milano, Dip.to di Meccanica, Via La Masa, 34, 20156 Milano, Italy Universidad de Oviedo, Dpto. Ciencia de los Materiales e Ingeniería Metalurgica, Campus de Gijón, Edificio de Energía, 33203 Gijón, Spain
a r t i c l e
i n f o
Article history: Received 13 September 2012 Received in revised form 15 February 2013 Accepted 19 March 2013 Available online 29 March 2013 Keywords: Ball valve Brittle fracture Notch effect Transition temperature
a b s t r a c t In this work the failure of a sub-sea ball valve, used in an oil-piping line, is analysed. The valve was of the same type and material already used for the construction of valves that were worked in service without any problem. The valve failed in the first pressure cycles during the preliminary laboratory tests, although the applied pressure was less than the design value. Metallographic and microstructural analysis of the fracture surfaces performed by means of optical and scanning electron microscope (SEM), residual stress and hardness measurement, tensile, toughness and Charpy tests, were executed in order to identify the causes of the failure. The results allowed assessing that the failure was due to two concomitant factors: a severe notch effect and an incorrect thermal treatment. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction The safe and reliable construction of mechanical systems depend both on the correct choice and the application of design methods aimed at avoiding severe stress conditions (commonly resulting in unexpected failures), as well as on the correct choice of the material and of the appropriate heat treatment, that should confer to the treated component the microstructural characteristics necessary to sustain the in service load conditions. For instance, the selected material should be able to resist the applied load and should be compatible with the surrounding environment that may include aggressive conditions, temperatures and other sources of damage [1]. The issue of guaranteeing the correct material properties is particularly important and insidious in the case of large components. In this case, the dimensions themselves can make the properties of the material quite variable in different zones. This fact involves the necessity of performing thermal treatments to obtain the desired properties. However it is quite unfeasible to gain uniform properties throughout the component geometry, especially due to the variable heating/cooling rates passing from the surface to the inner part of the components with large dimensions. If some machining operations are required after the thermal treatment, it is possible that local undesired properties appear in different zones. This can lead to critical stress conditions creating potential dangerous situations with the risk of catastrophic failures, from the economic point of view and sometime even in terms of human life loss. This remarks the importance of the correct, accurate and customized choice of the treatment parameters [2]. In this paper the failure analysis of a component with large dimensions is described. It is a 3000 subsea ball valve used in an oil piping line. The valve was of the same type and material already used for the construction of valves that were worked in service without any problem. The valve failed in the first pressure cycles during the preliminary laboratory tests, although the applied pressure was less than the in service one. The study includes different experimental observations and tests, which allow attributing the failure cause to two concomitant factors: a severe notch effect and incorrect thermal treatment, which lead to a brittle behaviour of the material. ⇑ Corresponding author. Tel.: +34 985 18 1992. E-mail addresses: [email protected] (S. Bagherifard), [email protected] (I. Fernández Pariente), [email protected] (M. Guagliano). 1350-6307/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.engfailanal.2013.03.012
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2. Case description The object of this study is a subsea ball valve used in oil piping line. The diameter of the ball is 3000 (762 mm). The expected pressure during usual operating conditions of the valve was equal to 50 MPa while the design pressure was set to 65 MPa. The in-service temperature of the valve was from 10 °C to 80 °C. The temperature during the test was 5 °C. The valve failed in the first pressure cycles during the preliminary laboratory tests, where only the internal pressure was the applied load. During the test the valve body was closed by a flanged joint closed by means of bolts and a gasket. The maximum value of pressure that was reached before the failure was 58 MPa. The material of the body of the valve is carbon steel type ASTM 694-F60. The geometry and a longitudinal section of the cited valve are shown in Fig. 1. The wall thickness of the valve body is equal to 350 mm. The entire inner surface of the valve body was cladded with Inconel 625 (tensile strength = 925 MPa, thickness of the clad 5 mm) to avoid corrosion problems. The steps carried out to determine the cause of the failure, are explained in detail in the sections: chemical and microstructural analysis, fractography (visual, optical, SEM), residual stresses and mechanical tests (tensile, hardness, toughness and Charpy tests).
3. Chemical and microstructural analysis Three sample series were cut from valve body just it the zone where the fracture took place. A series was cut adjacent to the internal diameter of the body (internal), the second one near the external diameter of the body (external) and the third in the middle of the wall thickness (middle). The analysis revealed that there were not significant differences in the chemical composition for different zones of the component. Table 1 shows the mean value of the spectrometric chemical analysis and the expected chemical composition according to the ASTM A694-F60 specification is also shown: no anomalies are noted. The specimens were metallographically prepared and chemically etched with a Nital 2% solution and then observed by optical microscope for analysing the microstructure of the material. Fig. 2a–c shows the microstructure of samples from different parts of the valve at different magnification factors (internal, medium and external part, respectively). The microstructure is regular and presents ferritic matrix with bainite, an acicular microstructure that forms in steels at temperatures of 250–550 °C depending on alloy content, and pearlite regularly distributed in the matrix. Metallurgical defects were not found. No remarkable microstructural differences were found considering the different analysed samples. As regards the grain size, finer grain dimensions are observed on the external zone, while in the internal region larger grains can be noted. This can be interpreted as a consequence of different cooling rates during the heat treatment [3].
4. Fractography 4.1. Visual fractography Fig. 3a shows a general view of the broken valve while Fig. 3b presents details of the fracture surface of the valve body. It is possible to see in Fig. 3c a probable point of crack initiation, and Fig. 3d shows a zone where the Inconel 625 cladding coat is lacking. In the visual observations it is not possible to identify plastic deformation. Brittle failure is intended as fracture that involves little or no plastic or permanent deformation [4]. Having this concept in mind the fracture surface evidences a brittle aspect that denotes unstable crack propagation. Beach marks and plastic deformation, which are characteristic of fatigue and
Fig. 1. (a) Valve body geometry and (b) longitudinal valve body section (the section was cut in the vertical plane of (a)).
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S. Bagherifard et al. / Engineering Failure Analysis 32 (2013) 167–177 Table 1 Spectrometric chemical analysis results. Element
C
Si
Mn
P
S
Cr
Ni
Mo
Al
%
0.182
0.21
1.11
0.004
0.001
0.12
0.15
0.40
0.022
Element
Cu
Co
Ti
Nb
V
B
Sn
As
N
%
0.12
0.0048
0.0017
0.0017
0.068
0.0002
0.0073
0.005
0.011
(Expected chemical composition: C 0.22 (max), Si 0.15–0.35, Mn 0.6–1.35, P 0.025 (max), S 0.025 (max), Ni 0.4 (max), CE = C + Mn/6+/Cr + Mo + V)/ 5 + (Ni + Cu/15) 0.42 (max).
Fig. 2. Microstructure of the valve: (a) internal part, (b) middle part, and (c) external part.
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(a)
(b)
(c)
(d)
Fig. 3. (a) General view of the broken valve (external diameter = 2 m), (b) Detail of the fracture surface, (c) crack initiation zone, and (d) detail of a zone not covered by the clad (by considering (a) as reference, (b) is at 3 o’clock, (c) at 1 o’clock and (d) at 6 o’clock).
ductile fracture respectively, were not noted. These considerations make possible to affirm that the failure was brittle and occurred in static condition. 4.2. Optical microscope fractography Optical microscope observations were also performed, by means of LEITZ ARISTOMET, on the three samples taken from the internal, external, and middle zone of the fracture surface (Fig. 4). These observations are aimed to evidence the morphology of the fracture surface and the fracture mechanism. It is to be mentioned that the fracture zone was marked before the fractography observations by a yellow1 colour coating, that is clearly noted in Fig. 4. These observations also evidence that the fracture surface analysis has a typically brittle fracture aspect. In some points a change of the fracture plane is observed as consequence of grain orientation, which is characteristic of brittle fracture. Nor fatigue striations and plastic deformation zones were observed. No remarkable differences between the three analysed zones (external, internal and middle) were found and no oxides changing the colour of the fracture surface were noted. This can be considered an confirmation of the instantaneous nature of the fracture, without pre-existent cracks. 4.3. SEM fractography The three samples cut from the internal, middle and external zones of the valve body were observed also by SEM using ZEISS EVO 50 microscope, to evidence the morphology of the fracture surface and the fracture mechanism. Fig. 5 reports some images taken from these different zones. The observation of these micrographs suggests that the failure was unstable, intergranular but with many transcrystalline facets and typically brittle. In all the analyzed three zones, the failure mechanism is cleavage without any observable plastic deformation. In some points it is possible to observe a change of the fracture plane as usually happens in brittle failures. No 1
For interpretation of color in Fig. 4, the reader is referred to the web version of this article.
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Fig. 4. Low magnification images (50) of (a) internal, (b) middle, and (c) external part of the valve (the pictures are taken at 1 o’clock of Fig. 3a).
fatigue striations were observed. The presence of inclusions was not detected in the investigated part of the fracture surface. No remarkable difference was observed between the three studied zones (external, internal and middle).
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Fig. 5. SEM micrograph of the valve (500) (a and b) internal part, (c and d) middle part, (e and f) external part (the zone where the pictures were taken is at 3 o’clock of Fig. 3a).
5. Residual stress measurement in the inner part of the body The measurement of the residual stresses on the internal Inconel 625 clad was executed by using an X-ray diffractometer (XRD) AST X-Stress 3000. Two points were chosen for performing measurements, one in the mid part of the body (see Fig. 6 to see the flange position), and the other at a distance of 100 mm from the flanged end of the body (see figure. It is to be underlined that the measurements were executed to have a general assessment of the effect of cladding in terms of residual stress state and that the measurements does not refer to the fracture zone since, due to the fracture itself it would have not been possible to get the residual stresses there. The area irradiated by X-ray is 2 2 mm2. The peak that was considered is positioned at 2h 148° The measurements were carried out using Cr Ka radiation considering the ‘‘w geometry’’ procedure. The diffraction peaks were detected by means of two PS detectors, symmetrically positioned with respect to the X-ray tube. The diffraction peak was measured in 11 different angular positions with regular intervals with respect of sin2 w = 0°, ±16.7°, ±24°, ±29.9°, ±35.1°, ±40°, and
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Fig. 6. The X-ray diffractometer placed on the inner part of the body of the valve.
Table 2 Results of the residual stresses measurements. Position
Middle 110 mm from the border
0°
45°
90°
r (MPa)
FWHM (°)
r (MPa)
FWHM (°)
r (MPa)
FWHM (°)
243 ± 50 195 ± 49
0.88 ± 0.17 1.44 ± 0,04
405 ± 105 161 ± 26
1.10 ± 0.02 1.49 ± 0,03
152 ± 55 –
0.95 ± 0.10 –
determined using the Cross Correlation method. For every considered angle the exposure time of the X-ray beam was equal to 50 s. A second result obtained from XRD measurement are the values of the width of the diffraction peaks at half of the peak height (FWHM), that can be consider an index of work hardening [5], were measured by XRD. The measurements were performed along three directions 0°, 90°, 45° that correspond, respectively to the axial direction of the valve, to the tangential one and to the one equi-spaced with respect of the previous ones. As regards the measurements at 100 mm from the flanged end of the valve body, a geometrical interference prevented the measurement at 90°, thus the calculation of the principal stresses was not possible. Fig. 6 shows the diffractometer taking measurements on the valve, and the measurement results are presented in Table 2. The obtained results indicate that the measured values are aligned with the ones usually expected from the technological process used for cladding, similar to welding and, consequently with the same order of magnitude of residual stresses. The variability of the measured values in the two measured points can be justified by considering the irregularity of the measured surface.
6. Mechanical tests 6.1. Tensile test The tensile tests were executed at room temperature according the ASTM 8 M [6]. Cylindrical specimens with a diameter of 12 mm, and an initial length of 60 mm were used. Tests were executed on 12 specimens taken from different parts of the valve, internal, external and in the middle of the thickness of the body, both, in axial and in tangential directions. All the tests were performed at room temperature under displacement control with a speed of 1 mm/min. Results do not denote remarkable differences on tested specimens, and show that none of the specimens meets the minimum requirement of the material in terms of UTS and yield stress. The mean values obtained from the test are resumed in Table 3. After the test, having a look at the broken surfaces, some inclusions were found in two of the tensile broken specimens; one of them is shown in Fig. 7. Those inclusions are related to the manufacturing process of the steel; indeed the presence of inclusions cannot be avoided but their dimension is related to the quality of the procedure and the process used for the steel manufacturing process. The appearance of the ‘‘fisheye’’ fracture shown in Fig. 7 could be also attributed to some form hydrogen embrittlement; anyway the chemical composition and the mechanical strength of the present steel are far from the ones typical of susceptibility to hydrogen and the valve was new and not used in environments leading to hydrogen embrittlement.
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Table 3 Resume of the results of the tensile tests. Young’s modulus (GPa)
Yield stress (Rp0.2) (MPa)
Peak stress (UTS) (MPa)
Elongation at break (%)
206.5
358
525
27.8
Expected minimum values, ASTM694-F60: Rp0.2 = 415, UTS = 515 MPa, Elongation at break 20%.
Fig. 7. Fish-eye failure starting from an internal inclusion happened in one of the tensile tests (41).
6.2. Hardness measurements In-depth micro hardness (HV) tests were performed on samples taken from the fracture surface according to the ASTM E-384 standard using a weight of 100 g and a duration of 15 s [7]. The measurements were performed on perpendicular sections to the fracture surface, on the internal, middle and external zones of the valve body, considering specimens cut from the zone at 3 o’clock in Fig. 3a. The micro-hardness trends are represented in Fig. 8. The micro-hardness values show a substantial uniformity along all the thickness of the body. The values are between 150HV and 200HV. 6.3. Toughness (KIC) tests The measurement of the fracture toughness was executed considering the method proposed by the ASTM 1820 code using an MTS machine 100 kN [8].The test were performed at room temperature (20 °C). Eighteen samples cut from different zones of the valve body were tested, considering longitudinal (L) and transversal (T) directions of cut with respect of the valve longitudinal axis. Three different zones were considered for cutting the samples: internal, middle and external zones of the valve body. The geometry is chosen according to ASTM E1820 code [8]. By elaborating the results it was possible to determine the critical value of J-Integral (JIC) and accordingly the KIC value could be assessed. In Table 4 the summary of the results got from 60 CTOD samples is reported.
200 180
HV
160 140 120 100 80
External Middle Internal
60 40 20 0
0
200
400
600
800
1000
1200
d (µm) Fig. 8. In-depth trend of micro-hardness along three different positions on the fracture surface (external, middle, internal).
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S. Bagherifard et al. / Engineering Failure Analysis 32 (2013) 167–177 Table 4 JIC and KIC values in samples from internal (I), external (E) and middle (M) part of the valve. E
JIC (N/mm) KIC (MPam0.5)
M
I
L
T
L
T
L
T
223 222
325 255
362 279
350 271
250 234
353 277
Table 5 Impact test (Charpy V-notch) results on samples from internal (I), external (E) and middle (M) part of the valve. I
KV8 (0 °C) (J) KV8 (RT °C) (J)
M
E
L
T
L
T
L
T
4.8 85.6
4.6 107.2
4.3 111.6
3.5 121.9
5.6 137.3
4.5 120.7
The results show a good average toughness value. KIC is higher than 180 MPam1/2 in all cases. However, results show a great variability even if specimens are taken from the same block (same zone and same orientation). It can be interpreted as an index of low quality of the result of the treatment or of the forging operation. Two specimens broke during the test, showing a fracture surface with brittle aspect: for these specimens it was not possible to determine the fracture toughness. Usually toughness values for steels vary from about 10 MPam1/2 to 250 MPam1/2 [9]. It can be concluded that the obtained experimental values are typical of a ductile material.
6.4. Charpy tests Also Charpy tests were executed on samples from internal, external and middle zone of the valve, cut in longitudinal and transversal orientation. Five samples of each zone and direction were tested (total number of tests = 30).Temperatures from 0 °C to room temperature with temperature tolerance of ±1 °C were considered, according to ASTM E23 for the Charpy V-Notch impact test [10]. The results in terms of the absorbed energy (KV8) are presented in Table 5. When the absorbed energy KV8 < 27 J, the material is considered brittle [11]. Very low values of toughness (smaller than 27 J) were measured on all the tests performed at 0 °C, that is that the material behaviour is strongly brittle at this temperature. The shear area was not measured on the fracture surfaces, since even by visual inspection it was clearly null and the fracture surface was completely flat. Both these items affirm the strong brittle behaviour of the material at 0 °C [12]. At room temperature the results were generally aligned even if they were not quite uniform. Some values strongly lower than the rest of the results are noted and they generally corresponded to the samples cut from the internal part of the valve body. In any case all the values are much bigger than 27 J and the measured shear areas were about 42%; this data evidences a relatively ductile behaviour at room temperature [11]. Having in count that the temperature recorded in the moment of the valve failure was around 0 °C, it is clear that fracture has taken place under brittle mechanism.
7. Discussion After having assessed the tests carried out on the broken valve, it is possible to affirm that there is not a unique well defined cause of the failure of the 3000 body of this valve. In fact, the experimental tests (tensile, CTOD, Charpy, hardness, residual stresses) and the material analysis (metallographic, microstructural, chemical) performed by different observation techniques (visual, SEM, optical microscope), did not evidence strong differences and anomalies. In particular, the metallurgical, microstructural and chemical analyses did not show unexpected results. Indeed, the behaviour of the material under tensile loads showed yields stress lower than the value prescribed by the ASTM A694/A694 M [12] for pipe flanges (316 vs. 415 MPa), while ultimate tensile stress is aligned with the minimum value required by the standards (522 vs. 515 MPa). The elongation measured to be 27%, is large enough to consider the material ductile and not brittle at room temperature. CTOD test show KIC values higher than 180 MPam1/2 in all cases. However, results represent a great variability, even if specimens are taken from the same block (same zone and same orientation). The poor uniformity of this data denotes the presence of some very local brittle zones that could have contributed to the failure of the valve.
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Considering the Charpy test results, the transition temperature of the valve material was estimated to be over 0 °C. Having in mind that the failure occurred at a temperature near 0 °C, it is concluded that the failure mechanism took place under brittle fracture conditions. It is also to be underlined that among the executed tests, some specimens showed anomalous behaviour: two tensile specimens broke due to the presence of large inclusions; two CTOD specimens broke during the test, showing brittle fracture surface, and finally some Charpy specimens resulted in relatively low values of the absorbed energy. On the whole, it can be said that the results of the Charpy tests present a great difference between tests carried out at room temperature and the ones performed at lower temperature, not only among the batch of specimens cut from the internal, middle or external zone but even among the specimens from the same zone: this poor uniformity can be considered as an index of not adequate quality of the material and can be related to the improper control of the heat treatment or the forging process. It is to be highlighted that the transition temperature is directly correlated to the forging conditions and can vary strongly in function of the parameters considered for the manufacturing process. It is also to be noted that the most anomalous results are obtained for the specimens cut from the internal part of the body. This fits well with the fact that failure originated from the inner part of the component body, also if these anomalies were not enough by themselves to justify the failure. The results of the fracture mechanics analysis exclude the possibility that the lack of penetration of the clad could be responsible of the failure. Indeed, the failure of the valve cannot be justified without considering the design of the valve body. In Fig. 1b the drawing of the section of the body with the failure location is shown. It can be noted that the failure happened in a zone of high stress concentration caused by the presence of a sharp notch with a small fillet radius. This particular geometry, under internal pressure induces a tri-axial stress state that can makes more brittle the behaviour of the material, thus, promote critical conditions [13] leading to a circumferential crack with respect of the point of maximum stress (the fillet with a small radius value). Hydrogen embrittlement could be considered as a possible cause of the failure; a tight dimensioning together with a small radius and some hydrogen embrittlement could be a possible cause of the failure. Anyway, steel ASTM A 694-F60 has both a chemical composition and a mechanical strength far from the typical ones that present this kind of damage. And the valve did not work in any environment able to justify an hydrogen embrittlement failure before. To come to the point, the executed analyses and the tests, suggest that the failure was principally due to two contemporary causes: the material behaviour, with marked variability not only within external, middle and internal body zones, but also within the same zone, and the particular design of the valve body, specifically in the zone of the failure section. A more uniform behaviour of the material could be obtained by strictly controlling the manufacturing process. In particular heat treatment could be optimized by performing some preliminary tests on pieces similar to the ones of interest and by controlling the real cooling rate in different zones of the valve. Having into account that forging parameters also have strong effects on the transition temperature, it is also important to pay special attention to the manufacturing conditions. From a design point of view, the sharp notch should be modified to obtain larger fillet radius. Also in this case, the execution of experimental tests under axial and pressure loads on cylindrical small scale specimens, that reproduce the geometrical characteristics of the critical part of the valve, could be useful in the design process. This kind of tests would allow determining whether under axial and pressure loads, the material behaviour tends to become brittle due to the tri-axial test state. 8. Conclusions Brittle failure mechanism was noted by analysing the broken valve body. The analysis proved that the material presented brittle behaviour at 0 °C, that it is to say that the transition temperature of the material is over 0 °C, which is not quite common for steels [14]. On the other hand, material toughness and absorbed energy in Charpy tests were found to be strongly variable not only in different zones of the valve body but also for the specimens taken from the same zone. The valve body was subjected to tensile stresses due to the applied pressure. Since the geometry of the valve presented in the failure zone a sharp notch a strong stress concentration factor and an almost equi-triaxial stress state were present in the crack initiation zone. In particular, the equi-triaxial stress state inhibits the plastic deformation (the von Mises stress is low) and the material behaviour tends to be brittle. Besides, in the moment of the failure, the temperature was around 0 °C, temperature at which the material showed a brittle behaviour. The concomitant appearance of these two factors (the notch effect and the high transition temperature) can explain the brittle failure [13]. Bearing in mind all these considerations, the conclusion is that the variable toughness, the high ductile to brittle transition temperature, and the load and stress conditions on the valve body made the material sensitive to the presence of a sharp notch, causing brittle fracture. References [1] Ashby MF. Materials Selection in Mechanical Design. Butterworth Heinemann; 2000. [2] Totten GE. Steel Heat Treatment. CRC Press Inc.; 2006.
S. Bagherifard et al. / Engineering Failure Analysis 32 (2013) 167–177 [3] [4] [5] [6] [7] [8] [9] [10] [11] [12]
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Handbook of Heat Treatment of Steels, 2008. D.J. Wulpi, Fractography Used to Analyse Fatigue and Brittle Fractures Features Conference ET 08, Orlando, USA, ASM, 2008. Noyan IC, Cohen JB. Residual Stress-Measurement by Diffraction and Interpretation. New York: Springer-Verlag; 1987. ASTM E8/E8M, Standard Test Methods for Tension Testing of Metallic Materials. ASTM E384, Standard Test Method for Knoop and Vickers Hardness of Materials. ASTM E1820, Standard Test Method for Measurement of Fracture Toughness. Asbhy M. Material Selection in Mechanical Design. Butterworth Heinemann; 2005. ASTM E23, Standard Test Methods for Notched Bar Impact Testing of Metallic Materials. ISO 148, Metallic Materials: Charpy Pendulum Impact Test. Part 1: Test Method. ASTM A694/ A 694M–08, Standard Specification for Carbon and Alloy Steel Forgings for Pipe Flanges, Fittings, Valves, and Parts for High-Pressure Transmission Service. [13] Wulpi DJ. Understanding How Components Fail. American Society for Metals; 1991. [14] Llewellyn DT, Hudd RC. Steels, Metallurgy and Applications. Butterworth-Heinemann; 2000.
Contents lists available at SciVerse ScienceDirect
Engineering Failure Analysis journal homepage: www.elsevier.com/locate/engfailanal
Failure analysis of a large ball valve for pipe-lines S. Bagherifard a, I. Fernández Pariente b,⇑, M. Guagliano a a b
Politecnico di Milano, Dip.to di Meccanica, Via La Masa, 34, 20156 Milano, Italy Universidad de Oviedo, Dpto. Ciencia de los Materiales e Ingeniería Metalurgica, Campus de Gijón, Edificio de Energía, 33203 Gijón, Spain
a r t i c l e
i n f o
Article history: Received 13 September 2012 Received in revised form 15 February 2013 Accepted 19 March 2013 Available online 29 March 2013 Keywords: Ball valve Brittle fracture Notch effect Transition temperature
a b s t r a c t In this work the failure of a sub-sea ball valve, used in an oil-piping line, is analysed. The valve was of the same type and material already used for the construction of valves that were worked in service without any problem. The valve failed in the first pressure cycles during the preliminary laboratory tests, although the applied pressure was less than the design value. Metallographic and microstructural analysis of the fracture surfaces performed by means of optical and scanning electron microscope (SEM), residual stress and hardness measurement, tensile, toughness and Charpy tests, were executed in order to identify the causes of the failure. The results allowed assessing that the failure was due to two concomitant factors: a severe notch effect and an incorrect thermal treatment. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction The safe and reliable construction of mechanical systems depend both on the correct choice and the application of design methods aimed at avoiding severe stress conditions (commonly resulting in unexpected failures), as well as on the correct choice of the material and of the appropriate heat treatment, that should confer to the treated component the microstructural characteristics necessary to sustain the in service load conditions. For instance, the selected material should be able to resist the applied load and should be compatible with the surrounding environment that may include aggressive conditions, temperatures and other sources of damage [1]. The issue of guaranteeing the correct material properties is particularly important and insidious in the case of large components. In this case, the dimensions themselves can make the properties of the material quite variable in different zones. This fact involves the necessity of performing thermal treatments to obtain the desired properties. However it is quite unfeasible to gain uniform properties throughout the component geometry, especially due to the variable heating/cooling rates passing from the surface to the inner part of the components with large dimensions. If some machining operations are required after the thermal treatment, it is possible that local undesired properties appear in different zones. This can lead to critical stress conditions creating potential dangerous situations with the risk of catastrophic failures, from the economic point of view and sometime even in terms of human life loss. This remarks the importance of the correct, accurate and customized choice of the treatment parameters [2]. In this paper the failure analysis of a component with large dimensions is described. It is a 3000 subsea ball valve used in an oil piping line. The valve was of the same type and material already used for the construction of valves that were worked in service without any problem. The valve failed in the first pressure cycles during the preliminary laboratory tests, although the applied pressure was less than the in service one. The study includes different experimental observations and tests, which allow attributing the failure cause to two concomitant factors: a severe notch effect and incorrect thermal treatment, which lead to a brittle behaviour of the material. ⇑ Corresponding author. Tel.: +34 985 18 1992. E-mail addresses: [email protected] (S. Bagherifard), [email protected] (I. Fernández Pariente), [email protected] (M. Guagliano). 1350-6307/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.engfailanal.2013.03.012
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2. Case description The object of this study is a subsea ball valve used in oil piping line. The diameter of the ball is 3000 (762 mm). The expected pressure during usual operating conditions of the valve was equal to 50 MPa while the design pressure was set to 65 MPa. The in-service temperature of the valve was from 10 °C to 80 °C. The temperature during the test was 5 °C. The valve failed in the first pressure cycles during the preliminary laboratory tests, where only the internal pressure was the applied load. During the test the valve body was closed by a flanged joint closed by means of bolts and a gasket. The maximum value of pressure that was reached before the failure was 58 MPa. The material of the body of the valve is carbon steel type ASTM 694-F60. The geometry and a longitudinal section of the cited valve are shown in Fig. 1. The wall thickness of the valve body is equal to 350 mm. The entire inner surface of the valve body was cladded with Inconel 625 (tensile strength = 925 MPa, thickness of the clad 5 mm) to avoid corrosion problems. The steps carried out to determine the cause of the failure, are explained in detail in the sections: chemical and microstructural analysis, fractography (visual, optical, SEM), residual stresses and mechanical tests (tensile, hardness, toughness and Charpy tests).
3. Chemical and microstructural analysis Three sample series were cut from valve body just it the zone where the fracture took place. A series was cut adjacent to the internal diameter of the body (internal), the second one near the external diameter of the body (external) and the third in the middle of the wall thickness (middle). The analysis revealed that there were not significant differences in the chemical composition for different zones of the component. Table 1 shows the mean value of the spectrometric chemical analysis and the expected chemical composition according to the ASTM A694-F60 specification is also shown: no anomalies are noted. The specimens were metallographically prepared and chemically etched with a Nital 2% solution and then observed by optical microscope for analysing the microstructure of the material. Fig. 2a–c shows the microstructure of samples from different parts of the valve at different magnification factors (internal, medium and external part, respectively). The microstructure is regular and presents ferritic matrix with bainite, an acicular microstructure that forms in steels at temperatures of 250–550 °C depending on alloy content, and pearlite regularly distributed in the matrix. Metallurgical defects were not found. No remarkable microstructural differences were found considering the different analysed samples. As regards the grain size, finer grain dimensions are observed on the external zone, while in the internal region larger grains can be noted. This can be interpreted as a consequence of different cooling rates during the heat treatment [3].
4. Fractography 4.1. Visual fractography Fig. 3a shows a general view of the broken valve while Fig. 3b presents details of the fracture surface of the valve body. It is possible to see in Fig. 3c a probable point of crack initiation, and Fig. 3d shows a zone where the Inconel 625 cladding coat is lacking. In the visual observations it is not possible to identify plastic deformation. Brittle failure is intended as fracture that involves little or no plastic or permanent deformation [4]. Having this concept in mind the fracture surface evidences a brittle aspect that denotes unstable crack propagation. Beach marks and plastic deformation, which are characteristic of fatigue and
Fig. 1. (a) Valve body geometry and (b) longitudinal valve body section (the section was cut in the vertical plane of (a)).
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S. Bagherifard et al. / Engineering Failure Analysis 32 (2013) 167–177 Table 1 Spectrometric chemical analysis results. Element
C
Si
Mn
P
S
Cr
Ni
Mo
Al
%
0.182
0.21
1.11
0.004
0.001
0.12
0.15
0.40
0.022
Element
Cu
Co
Ti
Nb
V
B
Sn
As
N
%
0.12
0.0048
0.0017
0.0017
0.068
0.0002
0.0073
0.005
0.011
(Expected chemical composition: C 0.22 (max), Si 0.15–0.35, Mn 0.6–1.35, P 0.025 (max), S 0.025 (max), Ni 0.4 (max), CE = C + Mn/6+/Cr + Mo + V)/ 5 + (Ni + Cu/15) 0.42 (max).
Fig. 2. Microstructure of the valve: (a) internal part, (b) middle part, and (c) external part.
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(a)
(b)
(c)
(d)
Fig. 3. (a) General view of the broken valve (external diameter = 2 m), (b) Detail of the fracture surface, (c) crack initiation zone, and (d) detail of a zone not covered by the clad (by considering (a) as reference, (b) is at 3 o’clock, (c) at 1 o’clock and (d) at 6 o’clock).
ductile fracture respectively, were not noted. These considerations make possible to affirm that the failure was brittle and occurred in static condition. 4.2. Optical microscope fractography Optical microscope observations were also performed, by means of LEITZ ARISTOMET, on the three samples taken from the internal, external, and middle zone of the fracture surface (Fig. 4). These observations are aimed to evidence the morphology of the fracture surface and the fracture mechanism. It is to be mentioned that the fracture zone was marked before the fractography observations by a yellow1 colour coating, that is clearly noted in Fig. 4. These observations also evidence that the fracture surface analysis has a typically brittle fracture aspect. In some points a change of the fracture plane is observed as consequence of grain orientation, which is characteristic of brittle fracture. Nor fatigue striations and plastic deformation zones were observed. No remarkable differences between the three analysed zones (external, internal and middle) were found and no oxides changing the colour of the fracture surface were noted. This can be considered an confirmation of the instantaneous nature of the fracture, without pre-existent cracks. 4.3. SEM fractography The three samples cut from the internal, middle and external zones of the valve body were observed also by SEM using ZEISS EVO 50 microscope, to evidence the morphology of the fracture surface and the fracture mechanism. Fig. 5 reports some images taken from these different zones. The observation of these micrographs suggests that the failure was unstable, intergranular but with many transcrystalline facets and typically brittle. In all the analyzed three zones, the failure mechanism is cleavage without any observable plastic deformation. In some points it is possible to observe a change of the fracture plane as usually happens in brittle failures. No 1
For interpretation of color in Fig. 4, the reader is referred to the web version of this article.
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Fig. 4. Low magnification images (50) of (a) internal, (b) middle, and (c) external part of the valve (the pictures are taken at 1 o’clock of Fig. 3a).
fatigue striations were observed. The presence of inclusions was not detected in the investigated part of the fracture surface. No remarkable difference was observed between the three studied zones (external, internal and middle).
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Fig. 5. SEM micrograph of the valve (500) (a and b) internal part, (c and d) middle part, (e and f) external part (the zone where the pictures were taken is at 3 o’clock of Fig. 3a).
5. Residual stress measurement in the inner part of the body The measurement of the residual stresses on the internal Inconel 625 clad was executed by using an X-ray diffractometer (XRD) AST X-Stress 3000. Two points were chosen for performing measurements, one in the mid part of the body (see Fig. 6 to see the flange position), and the other at a distance of 100 mm from the flanged end of the body (see figure. It is to be underlined that the measurements were executed to have a general assessment of the effect of cladding in terms of residual stress state and that the measurements does not refer to the fracture zone since, due to the fracture itself it would have not been possible to get the residual stresses there. The area irradiated by X-ray is 2 2 mm2. The peak that was considered is positioned at 2h 148° The measurements were carried out using Cr Ka radiation considering the ‘‘w geometry’’ procedure. The diffraction peaks were detected by means of two PS detectors, symmetrically positioned with respect to the X-ray tube. The diffraction peak was measured in 11 different angular positions with regular intervals with respect of sin2 w = 0°, ±16.7°, ±24°, ±29.9°, ±35.1°, ±40°, and
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Fig. 6. The X-ray diffractometer placed on the inner part of the body of the valve.
Table 2 Results of the residual stresses measurements. Position
Middle 110 mm from the border
0°
45°
90°
r (MPa)
FWHM (°)
r (MPa)
FWHM (°)
r (MPa)
FWHM (°)
243 ± 50 195 ± 49
0.88 ± 0.17 1.44 ± 0,04
405 ± 105 161 ± 26
1.10 ± 0.02 1.49 ± 0,03
152 ± 55 –
0.95 ± 0.10 –
determined using the Cross Correlation method. For every considered angle the exposure time of the X-ray beam was equal to 50 s. A second result obtained from XRD measurement are the values of the width of the diffraction peaks at half of the peak height (FWHM), that can be consider an index of work hardening [5], were measured by XRD. The measurements were performed along three directions 0°, 90°, 45° that correspond, respectively to the axial direction of the valve, to the tangential one and to the one equi-spaced with respect of the previous ones. As regards the measurements at 100 mm from the flanged end of the valve body, a geometrical interference prevented the measurement at 90°, thus the calculation of the principal stresses was not possible. Fig. 6 shows the diffractometer taking measurements on the valve, and the measurement results are presented in Table 2. The obtained results indicate that the measured values are aligned with the ones usually expected from the technological process used for cladding, similar to welding and, consequently with the same order of magnitude of residual stresses. The variability of the measured values in the two measured points can be justified by considering the irregularity of the measured surface.
6. Mechanical tests 6.1. Tensile test The tensile tests were executed at room temperature according the ASTM 8 M [6]. Cylindrical specimens with a diameter of 12 mm, and an initial length of 60 mm were used. Tests were executed on 12 specimens taken from different parts of the valve, internal, external and in the middle of the thickness of the body, both, in axial and in tangential directions. All the tests were performed at room temperature under displacement control with a speed of 1 mm/min. Results do not denote remarkable differences on tested specimens, and show that none of the specimens meets the minimum requirement of the material in terms of UTS and yield stress. The mean values obtained from the test are resumed in Table 3. After the test, having a look at the broken surfaces, some inclusions were found in two of the tensile broken specimens; one of them is shown in Fig. 7. Those inclusions are related to the manufacturing process of the steel; indeed the presence of inclusions cannot be avoided but their dimension is related to the quality of the procedure and the process used for the steel manufacturing process. The appearance of the ‘‘fisheye’’ fracture shown in Fig. 7 could be also attributed to some form hydrogen embrittlement; anyway the chemical composition and the mechanical strength of the present steel are far from the ones typical of susceptibility to hydrogen and the valve was new and not used in environments leading to hydrogen embrittlement.
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Table 3 Resume of the results of the tensile tests. Young’s modulus (GPa)
Yield stress (Rp0.2) (MPa)
Peak stress (UTS) (MPa)
Elongation at break (%)
206.5
358
525
27.8
Expected minimum values, ASTM694-F60: Rp0.2 = 415, UTS = 515 MPa, Elongation at break 20%.
Fig. 7. Fish-eye failure starting from an internal inclusion happened in one of the tensile tests (41).
6.2. Hardness measurements In-depth micro hardness (HV) tests were performed on samples taken from the fracture surface according to the ASTM E-384 standard using a weight of 100 g and a duration of 15 s [7]. The measurements were performed on perpendicular sections to the fracture surface, on the internal, middle and external zones of the valve body, considering specimens cut from the zone at 3 o’clock in Fig. 3a. The micro-hardness trends are represented in Fig. 8. The micro-hardness values show a substantial uniformity along all the thickness of the body. The values are between 150HV and 200HV. 6.3. Toughness (KIC) tests The measurement of the fracture toughness was executed considering the method proposed by the ASTM 1820 code using an MTS machine 100 kN [8].The test were performed at room temperature (20 °C). Eighteen samples cut from different zones of the valve body were tested, considering longitudinal (L) and transversal (T) directions of cut with respect of the valve longitudinal axis. Three different zones were considered for cutting the samples: internal, middle and external zones of the valve body. The geometry is chosen according to ASTM E1820 code [8]. By elaborating the results it was possible to determine the critical value of J-Integral (JIC) and accordingly the KIC value could be assessed. In Table 4 the summary of the results got from 60 CTOD samples is reported.
200 180
HV
160 140 120 100 80
External Middle Internal
60 40 20 0
0
200
400
600
800
1000
1200
d (µm) Fig. 8. In-depth trend of micro-hardness along three different positions on the fracture surface (external, middle, internal).
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S. Bagherifard et al. / Engineering Failure Analysis 32 (2013) 167–177 Table 4 JIC and KIC values in samples from internal (I), external (E) and middle (M) part of the valve. E
JIC (N/mm) KIC (MPam0.5)
M
I
L
T
L
T
L
T
223 222
325 255
362 279
350 271
250 234
353 277
Table 5 Impact test (Charpy V-notch) results on samples from internal (I), external (E) and middle (M) part of the valve. I
KV8 (0 °C) (J) KV8 (RT °C) (J)
M
E
L
T
L
T
L
T
4.8 85.6
4.6 107.2
4.3 111.6
3.5 121.9
5.6 137.3
4.5 120.7
The results show a good average toughness value. KIC is higher than 180 MPam1/2 in all cases. However, results show a great variability even if specimens are taken from the same block (same zone and same orientation). It can be interpreted as an index of low quality of the result of the treatment or of the forging operation. Two specimens broke during the test, showing a fracture surface with brittle aspect: for these specimens it was not possible to determine the fracture toughness. Usually toughness values for steels vary from about 10 MPam1/2 to 250 MPam1/2 [9]. It can be concluded that the obtained experimental values are typical of a ductile material.
6.4. Charpy tests Also Charpy tests were executed on samples from internal, external and middle zone of the valve, cut in longitudinal and transversal orientation. Five samples of each zone and direction were tested (total number of tests = 30).Temperatures from 0 °C to room temperature with temperature tolerance of ±1 °C were considered, according to ASTM E23 for the Charpy V-Notch impact test [10]. The results in terms of the absorbed energy (KV8) are presented in Table 5. When the absorbed energy KV8 < 27 J, the material is considered brittle [11]. Very low values of toughness (smaller than 27 J) were measured on all the tests performed at 0 °C, that is that the material behaviour is strongly brittle at this temperature. The shear area was not measured on the fracture surfaces, since even by visual inspection it was clearly null and the fracture surface was completely flat. Both these items affirm the strong brittle behaviour of the material at 0 °C [12]. At room temperature the results were generally aligned even if they were not quite uniform. Some values strongly lower than the rest of the results are noted and they generally corresponded to the samples cut from the internal part of the valve body. In any case all the values are much bigger than 27 J and the measured shear areas were about 42%; this data evidences a relatively ductile behaviour at room temperature [11]. Having in count that the temperature recorded in the moment of the valve failure was around 0 °C, it is clear that fracture has taken place under brittle mechanism.
7. Discussion After having assessed the tests carried out on the broken valve, it is possible to affirm that there is not a unique well defined cause of the failure of the 3000 body of this valve. In fact, the experimental tests (tensile, CTOD, Charpy, hardness, residual stresses) and the material analysis (metallographic, microstructural, chemical) performed by different observation techniques (visual, SEM, optical microscope), did not evidence strong differences and anomalies. In particular, the metallurgical, microstructural and chemical analyses did not show unexpected results. Indeed, the behaviour of the material under tensile loads showed yields stress lower than the value prescribed by the ASTM A694/A694 M [12] for pipe flanges (316 vs. 415 MPa), while ultimate tensile stress is aligned with the minimum value required by the standards (522 vs. 515 MPa). The elongation measured to be 27%, is large enough to consider the material ductile and not brittle at room temperature. CTOD test show KIC values higher than 180 MPam1/2 in all cases. However, results represent a great variability, even if specimens are taken from the same block (same zone and same orientation). The poor uniformity of this data denotes the presence of some very local brittle zones that could have contributed to the failure of the valve.
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Considering the Charpy test results, the transition temperature of the valve material was estimated to be over 0 °C. Having in mind that the failure occurred at a temperature near 0 °C, it is concluded that the failure mechanism took place under brittle fracture conditions. It is also to be underlined that among the executed tests, some specimens showed anomalous behaviour: two tensile specimens broke due to the presence of large inclusions; two CTOD specimens broke during the test, showing brittle fracture surface, and finally some Charpy specimens resulted in relatively low values of the absorbed energy. On the whole, it can be said that the results of the Charpy tests present a great difference between tests carried out at room temperature and the ones performed at lower temperature, not only among the batch of specimens cut from the internal, middle or external zone but even among the specimens from the same zone: this poor uniformity can be considered as an index of not adequate quality of the material and can be related to the improper control of the heat treatment or the forging process. It is to be highlighted that the transition temperature is directly correlated to the forging conditions and can vary strongly in function of the parameters considered for the manufacturing process. It is also to be noted that the most anomalous results are obtained for the specimens cut from the internal part of the body. This fits well with the fact that failure originated from the inner part of the component body, also if these anomalies were not enough by themselves to justify the failure. The results of the fracture mechanics analysis exclude the possibility that the lack of penetration of the clad could be responsible of the failure. Indeed, the failure of the valve cannot be justified without considering the design of the valve body. In Fig. 1b the drawing of the section of the body with the failure location is shown. It can be noted that the failure happened in a zone of high stress concentration caused by the presence of a sharp notch with a small fillet radius. This particular geometry, under internal pressure induces a tri-axial stress state that can makes more brittle the behaviour of the material, thus, promote critical conditions [13] leading to a circumferential crack with respect of the point of maximum stress (the fillet with a small radius value). Hydrogen embrittlement could be considered as a possible cause of the failure; a tight dimensioning together with a small radius and some hydrogen embrittlement could be a possible cause of the failure. Anyway, steel ASTM A 694-F60 has both a chemical composition and a mechanical strength far from the typical ones that present this kind of damage. And the valve did not work in any environment able to justify an hydrogen embrittlement failure before. To come to the point, the executed analyses and the tests, suggest that the failure was principally due to two contemporary causes: the material behaviour, with marked variability not only within external, middle and internal body zones, but also within the same zone, and the particular design of the valve body, specifically in the zone of the failure section. A more uniform behaviour of the material could be obtained by strictly controlling the manufacturing process. In particular heat treatment could be optimized by performing some preliminary tests on pieces similar to the ones of interest and by controlling the real cooling rate in different zones of the valve. Having into account that forging parameters also have strong effects on the transition temperature, it is also important to pay special attention to the manufacturing conditions. From a design point of view, the sharp notch should be modified to obtain larger fillet radius. Also in this case, the execution of experimental tests under axial and pressure loads on cylindrical small scale specimens, that reproduce the geometrical characteristics of the critical part of the valve, could be useful in the design process. This kind of tests would allow determining whether under axial and pressure loads, the material behaviour tends to become brittle due to the tri-axial test state. 8. Conclusions Brittle failure mechanism was noted by analysing the broken valve body. The analysis proved that the material presented brittle behaviour at 0 °C, that it is to say that the transition temperature of the material is over 0 °C, which is not quite common for steels [14]. On the other hand, material toughness and absorbed energy in Charpy tests were found to be strongly variable not only in different zones of the valve body but also for the specimens taken from the same zone. The valve body was subjected to tensile stresses due to the applied pressure. Since the geometry of the valve presented in the failure zone a sharp notch a strong stress concentration factor and an almost equi-triaxial stress state were present in the crack initiation zone. In particular, the equi-triaxial stress state inhibits the plastic deformation (the von Mises stress is low) and the material behaviour tends to be brittle. Besides, in the moment of the failure, the temperature was around 0 °C, temperature at which the material showed a brittle behaviour. The concomitant appearance of these two factors (the notch effect and the high transition temperature) can explain the brittle failure [13]. Bearing in mind all these considerations, the conclusion is that the variable toughness, the high ductile to brittle transition temperature, and the load and stress conditions on the valve body made the material sensitive to the presence of a sharp notch, causing brittle fracture. References [1] Ashby MF. Materials Selection in Mechanical Design. Butterworth Heinemann; 2000. [2] Totten GE. Steel Heat Treatment. CRC Press Inc.; 2006.
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Handbook of Heat Treatment of Steels, 2008. D.J. Wulpi, Fractography Used to Analyse Fatigue and Brittle Fractures Features Conference ET 08, Orlando, USA, ASM, 2008. Noyan IC, Cohen JB. Residual Stress-Measurement by Diffraction and Interpretation. New York: Springer-Verlag; 1987. ASTM E8/E8M, Standard Test Methods for Tension Testing of Metallic Materials. ASTM E384, Standard Test Method for Knoop and Vickers Hardness of Materials. ASTM E1820, Standard Test Method for Measurement of Fracture Toughness. Asbhy M. Material Selection in Mechanical Design. Butterworth Heinemann; 2005. ASTM E23, Standard Test Methods for Notched Bar Impact Testing of Metallic Materials. ISO 148, Metallic Materials: Charpy Pendulum Impact Test. Part 1: Test Method. ASTM A694/ A 694M–08, Standard Specification for Carbon and Alloy Steel Forgings for Pipe Flanges, Fittings, Valves, and Parts for High-Pressure Transmission Service. [13] Wulpi DJ. Understanding How Components Fail. American Society for Metals; 1991. [14] Llewellyn DT, Hudd RC. Steels, Metallurgy and Applications. Butterworth-Heinemann; 2000.