Assessment of hot tears in cast steel components

International Journal of Pressure Vessels and Piping 78 (2001) 865±874

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Assessment of hot tears in cast steel components K. Maile a,*, H. Theofel a, C. Weichert a, K.-H. Mayer b, C. Gerdes c, S. Sheng d a

Staatliche MaterialpruÈfungsanstalt (MPA), University of Stuttgart, Stuttgart, Germany b Alstom Power, NuÈrnberg, Germany c Alstom Power Ltd, Baden, Switzerland d SIEMENS AG, Power Generation Group (KWU), MuÈlheim/Ruhr, Germany

Abstract Large scale specimens with real internal defects were taken from castings. The behaviour of the defects under service like loading conditions (creep, creep±fatigue and fatigue tests at 5308C) was studied, accompanied by non-destructive testing and post-test investigations. q 2002 Elsevier Science Ltd. All rights reserved. Keywords: Creep resistant cast steels; Natural defects; Hot tears

1. Introduction Creep resistant cast steels are used for high temperature components of power stations and other plants, in order to combine improved static and dynamic strength and economic manufacturing. The current loading of components in modern power plants is characterised by increased cyclic service conditions. Consequently, the optimisation of component geometry in combination with the improvement of the casting technique, the improvement of the inspection technique, (including the realistic evaluation of non destructive ®ndings), as well as knowledge of material dependent damage mechanics is required. Due to physical and chemical processes during the solidi®cation, typical casting defects can arise, such as oxide ®lm, grain boundary cracks, moulding material inclusions and volume defects, (micropores, shrink holes and hot tears). These natural defects, in particular hot tears or healed hot tears with segregation zones, occur despite of the improvement of the manufacturing quality. If of critical size and in critical places of higher stressed, areas they can lead to failure due to crack growth. Therefore, in addition to the improvement of non destructive testing methods with regard to the reliable determination of size, position and distribution of defects in components, it is necessary to create methods to evaluate the acceptance of potential manufacturing ¯aws in the castings with speci®c consideration of the actual loading situation [1,2]. This also relates to the question how natural * Corresponding author. E-mail address: [email protected] (K. Maile).

defects behave concerning crack initiation and crack propagation. For the description and evaluation of this behaviour, the following parameters, schematically represented in Fig. 1, have to be considered: ² reliable detection of the defect formation, with regard to size, position and distribution, by non destructive testing (NDT); ² global and local stresses, taking into account the magnitude, duration and variation with time; ² material properties in the vicinity of a natural defect in comparison to the defect-free material; ² in¯uence of surface or near surface position of defects and mutual interaction of defects; ² the simpli®cation, that the natural defects behave like cracks; ² the idealisation of irregularly formed defects as circles and ellipses. A better knowledge for a quantitative description of a possible crack initiation or crack propagation at tolerated defects leads to a more economical operation by higher utilisation of material or by longer assessment intervals. 2. Formation and characterisation of hot tears The types of surface defects and internal defects are documented and described in Ref. [3]. Besides shrink holes, hot tears are of major in¯uence on the component integrity. The formation and development of these ¯aws near the surface is described in Ref. [4]. In general hot

0308-0161/02/$ - see front matter q 2002 Elsevier Science Ltd. All rights reserved. PII: S 0308-016 1(01)00101-6

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Fig. 1. Variables for the evaluation of natural defects.

tears develop during the solidi®cation process, just above or at the solidus temperature. The main reason for their formation are problems with the shaping of the cast piece and a restrained shrinking process, caused by the geometry. Local stresses in the component lead to intergranular separations with complex 3D-shapes. In the case of discontinuous solidi®cation, enrichment of certain elements takes place in the residual melt, leading to an increased solidi®cation temperature and extension of the solidi®cation time, and promoting the development of additional hot tears. Generally hot tears do not have a connection to the casting surface. Due to their formation at temperatures within the solidi®cation interval, the internal surface of the defects shows a typical solidi®cation structure and the shape of the ¯aws is rounded. A combination of macro segregations and micro shrink holes can often be observed. Beside open hot tears, partly healed and completely healed hot tears can also be found. They result from penetration of residual melt, which ®lls up the separations totally (healed hot tears) or partly (partly healed hot tears). 3. NDT-methods For indication and evaluation of defects and cracks at surfaces as well as in the volume of components, different NDT procedures are applied. For the detection of internal defects, situated more or less close to the surface, ultrasonic test methods (UT) worked satisfactorily. To indicate open and partly healed hot tears and surface cracks, the magnetic particle test (MP) is most suitable. Beside the main UT and MP procedures the X-ray inspection, dye penetration test, eddy current inspection, and acoustic emission analysis are also used. The evaluation of detected defects in components is performed according to national standards or company speci®cations. Thus, the examination and de®nition of resolution and allowable limits are set up according to Ref. [5], whereas the surface inspection takes place according to Ref. [6]. For example for UT inspection, coherent

defects with distinct increase in length and depth must be registered, regardless of their echo height. Also in a new European standard [7] it is considered, that echo heights are not relevant, and only the defect increase indicated with UT as well as the type of the echo dynamics formation is important. Resolution and allowable limits depend, apart from the quality grade, on the depth position of the defect. For example, defects are not acceptable for the quality grade 1 but acceptable for the quality grades 2±4, if their presence is allowed by a fracture-mechanics evaluation. If the critical defect size is exceeded, the component must be rejected or the detected defects have to be eliminated by a repair measure. For indications near the surface, the echo height is important as an evaluation criterion. In practice, due to the high standard of NDT, hot tears and partly healed hot tears can be reliably detected [8,9]. Progress in the determination of defect sizes could be achieved on the basis of echo-dynamic procedure [10]. Modern computerassisted UT methods for the determination of the position and growth of indications have improved the description of defects and the reproducibility of their detecting [11]. The typical echo dynamics of sub-surface indications in steel castings is shown schematically in Fig. 2. As well as defects near the surface, the partly healed hot tears (type of UT display NIP) are of special interest. For these group indications in particular, the non destructive determination of the distance between defects is dif®cult. First hints about the average defect distance in resolvable group indications in forgings could be derived from the number of echo peaks as well as from the analysis of the wave attenuation [12]. The evaluation of group indications in standards, e.g. Refs. [13,14], is simpli®ed and made by summation to one individual defect. This may sometimes lead to an extraordinarily conservative evaluation. 4. Experimental investigations As test material, cast rings of type 1CrMoV (grade

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Fig. 2. Characteristic display forms at UT checks with angle probes.

GS-17CrMoV5-11), Fig. 3 were used. These samples were produced as pilot castings for these investigations used various castings techniques to obtain internal defects, mainly hot tears, in the transition radius to the thick walled ¯ange. From this region, based on NDT-indications, large scale specimens were removed, with natural defects in the middle of the specimen length. The specimens had a square cross section 40 £ 40 mm, with two concave surfaces, one of them as the original cast surface and two parallel plane surfaces. The procedure for the removal, testing and posttest investigations is represented in Fig. 4. The removal and testing of the feature test specimens was monitored by UT: ∅ 1480 ∅ 884

∅ 1095 R 200 X-Maß 150 ∅ 1820 Fig. 3. 1CrMoV cast ring, from which the large scale specimens were removed.

after a ®rst inspection at the component (US-I) another check was carried out at the test bars (US-II). Based on these ®ndings, the specimens were removed and manufactured in such a way that natural defects were embedded in the shaft of the sample. Further UT was carried out after the manufacturing of specimens (US-III), as well as during and after the loading (US-IV), which comprised creep, creep±fatigue and fatigue tests, conducted at 5308C. The specimens were then broken under liquid nitrogen, to make the internal defects accessible for post test investigations. The fracture surface especially in the vicinity of the defects was investigated by means of a scanning electron microscope, SEM, in order to detect both damage and crack growth, as well as the type of defects. The real defect size found by SEM were compared with the UT results. During the test, the elongation in the specimen center was measured by means of capacitive gauges and clip gauges with a measuring length of 20 mm. Also, the AC eddy current probe was used to detect possible crack initiation and propagation. Fig. 5 shows results of some representative tests under creep fatigue loading. The static portion in these tests was dominating, due to holding times of 20±720 min at maximum load. The creep curves for these tests show the typical behaviour. Specimens with higher net stress and specimens with larger defects …s n @ s brutto ; whereas s brutto ˆ F=Atotal ) yield shortest test durations. The long term tests have been terminated on reaching 1% plastic strain in accordance with the deformation limits used in service. Also former investigations showed, that in this range, up to 1% plastic strain deformation induced transgranular cracking occurs [15].

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σ number of cycles

σ σ

number of cycles

t (kh)

Fig. 4. Large scale specimens from castings with internal defects, schematic representation of specimen removal, loading and investigations.

In the following, two tests (s brutto ˆ 181 MPa; 1472 h and s brutto ˆ 152 MPa; 10351 h) will be used to demonstrate speci®cally: ² the behaviour of hot tears in terms of damage mechanism taking into account the speci®c loading situation using SEM investigations, ² the accuracy of UT by comparing the SEM ®ndings with UT results with respect to size and location, ² the possibility to using fracture mechanics to evaluate the allowable size of defects with regard to crack initiation and propagation, based on the standard elastic fracture mechanic parameter K, and calculations using the Finite Elemente Method (FEM). 2.0

small sc. specimen 180 MPa

4.1. SEM-Investigations The specimens have been fractured, in liquid nitrogen, after completion of the tests. Thus the hot tears are embedded in a fracture surface with typical brittle character and the identi®cation of the respective zones (hot tears, crack propagation) could be done. Fig. 6 shows the macroscopic view of the fracture surface, where the initial hot tears have been marked. During the test, crack initiation and propagation through regions of healed hot tears has occurred and the different hot tears have grown together. Figs. 7±9 show details from the fracture surface at higher magni®cation. The typical structure of the solidi®ed surface of both an open hot tear t = 20 min h th = 20 min th = 60 min t = 60 min h th = 60 min t = 720 min h th = 720 min small scale sp.

small sc. specimen 155 MPa

σbrutto=181 MPa

pl. strain (%)

1.5

180

1.0

152 152

#

#

155

163

0.5

158

# = test interruption

0.0 0

2000

4000 6000 8000 time at max. load (h)

10000

12000

Fig. 5. Creep fatigue tests at large scale specimens, course of the plastic strain versus time.

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4.2. Evaluation of the UT-signals A

B C

5 mm

×

cast surface ×

Fig. 6. Creep±fatigue specimen, t ˆ 1472 h; part of forced fracture surface, overview.

as well as a healed hot tear is visible. Also the typical intergranular creep crack growth at the transition from the healed hot tear to the base material can be seen, Fig. 9.

Fig. 10 shows the results of the C-scan (specimen s brutto ˆ 181 MPa; 1472 h) before loading and at the end of the test. The location was at a distance of 166 mm from the specimen's end face and it can be seen that a distinctive increase of the signals has taken place. The real sizes, Fig. 6, ®ts well to the indicated sizes before and after loading. Also there is good agreement with location in the specimen. A summary of the ultrasonic ®ndings in samples prior to the start of the test as compared to the depths and lengths of ¯aws discovered at the brittle fractured areas is shown in Fig. 11. The maximum ¯aw depths shown on the left-hand site show very good agreement for about 60% of the ¯aws. The other ¯aw depths determined by ultrasonic methods are conservative. The ¯aw length comparison shown on the right-hand graph exhibit even better agreement [15]. 4.3. Fracture mechanics and FEM calculation To calculate the local time dependent stress at the defects, the program abaqus was used. Fig. 12 shows the FE-mesh for the specimen (creep±fatigue-test, 10351 h) and the defects at the corner of the specimen have been modelled

Detail A from fig. 6 hot tear with free soldified surface and with inclusions

50 µm

Fig. 7. Detail A from Fig. 6 hot tear with free solidi®ed surface and with inclusions.

Detail B from fig. 6 opened healed hot tear

50 µm

Fig. 8. Detail B from Fig. 6 opened healed hot tear.

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Detail C from fig. 6 healed hot tear crack growth in base metal and cleavage fracture ( ∆a ≈ 150 µm)

50 µm

Fig. 9. Detail C from Fig. 6 healed hot tear crack growth in base metal and cleavage fracture (! Da < 150 mm).

before loading

after 1472 h

Fig. 10. Creep±fatigue specimen, t ˆ 1472 h; C-scans from the region x ˆ 166 mm (distance from specimen end face). 40

20 internal up to surface at corner long up to surface

III u. IV V

15

max. Length of Flaws in mm

max. Depth of Flaws in mm

I u. II

IV

Location

Flaw Typ

Location

Flaw Typ

10

5

I u. II

35

III u. IV V

30

IV

internal up to surface at corner long up to surface

25 20 15 10 5 0

0 0

5

10

15

20

Depth of Flaws determinated by US in mm

0

5

10

15

20

25

30

35

40

Length of Flaws determinated by US in mm

Fig. 11. Comparison of true ¯aw size with the results of UT inspection of large specimens (GS-17CrMoV5-11).

on basis of the fractographic investigation. For the model a sharp crack was used. The crack tip radius was assumed zero. Due to plastic deformation during loading, no differences to the results of a calculation with a small crack tip radius is expected. A Norton±Bailey type creep law was used for time dependent calculations. The results of UT are shown in Fig. 13 prior to, during and after loading.

This indicates that crack initiation and signi®cant propagation took place after at least 3500 h. The time-dependant evolution of the von Mises stress shows signi®cant stress redistribution and relaxation in the vicinity of defects whereas the net stress in the ligament is constant. The net stress calculated by FEM in axial direction is in good accordance with analytically calculated axial stress. The

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half large scale specimen

3 2

1

casting surface Defect A

node 1089

Defect B node 880

node 4534

node 5002

detail from corner with defects

3 2

1

node 6658 node 7353

Defect C

Fig. 12. FE-mesh for large scale specimen with defects near corner (creep±fatigue test, 10351 h).

before loading and after 3495 h

after 10351 h

Fig. 13. Creep±fatigue specimen, t ˆ 10351 h; C-scans from the region x ˆ 192±200 mm (distance from specimen end face).

comparison of the values of stress intensity factor K, obtained by FEM and from analytical formulae show a good coincidence if the defects were modelled with an elliptical shape …KI analytical ˆ 547 N=mm3=2 ; KI FE ˆ 516 N=mm3=2 †: The FEM analysis, Fig. 14, gives lower values for the single defects, for example, KI FE ˆ 350 N=mm3=2 ; defect B in Fig. 12. 4.4. Evaluation by the two-criteria-diagram In Fig. 15 the crack initiation values for GS-17CrMoV511 are shown in a two-criteria-diagram, which was proposed

by Ewald [16]. Results obtained using compact tension specimens (Cs25) and double edge notched tensile specimens of different sizes (DENT with a width of 9,15, 30 and 60 mm). Results of longtime tests on cast steel components with manufacturing defects are also included. A more detailed description of the practical use of this method is given in Ref. [17]. The boundary lines de®ning crack initiation are validated by these results. Therefore the two-criteria-diagram is practicable to determine the crack initiation time for large scale specimens with defects. To this aim, the damage paths of a specimen is plotted in the diagram. Fig. 16 shows damage paths for three natural

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3/2

3/2 Spannungsintesitätfactor [N/mm stress intensity factor [N/mm ] ]

872

prior to loading, was used for the evaluation of crack initiation and propagation. In parallel, the net stress concept was used to evaluate crack initiation at large scale specimens with embedded defects. The basis for this evaluation were results of defect free small scale specimens under pure creep loading. In Fig. 17 the net stress Ð Rp0,2 ratio is plotted over time for crack initiation, which was determined using the two-criteria-diagram (open symbols). Closed symbols show the same ratio for the end of the test. The creep rupture strength for small scale specimens, and the 1% creep strain curve for the defect free creep specimens is also shown. The lower scatterband for cracks below resolution limit and cracks which are not acceptable is also depicted. It can be seen, that the crack initiation time determined by the two-criteria-diagram is conservative. The Rp1 curve of defect free creep specimens can be taken as a lower bound to assess the crack initiation behaviour for specimens having defects. In Fig. 18 the K-parameter is used to evaluate the crack initiation time at defects. The dashed base line was obtained from creep crack growth tests with defect free compact tension specimen, with a width of 25 mm (CT25). For all specimens, mixed-mode or ligament damage can be determined, this means, that damage is due to creep in the ligament. The elastic fracture mechanics parameter K for the specimens with defects at crack initiation and the end of the test are below the CT25 crack initiation curve.

SFG220, defect Riß Nr.1 B 2000

1000

node 1089 Knoten 1089 node 880 Knoten 880

0

5

10

Abstand [mm]

distance [mm]

Fig. 14. Stress intensity factor versus the contour of defect B.

defects of the specimen SFG220 (s brutto ˆ 152 MPa; 10351 h). The point of intersection between the path and the boundary line delivers the theoretical point of crack initiation. The paths of this and the other examined specimens show ligament damage or mixed mode damage as failure mode. 4.5. Evaluation by means of linear elastic fracture mechanics The linear elastic fracture mechanics parameter K was used to describe the stress situation at the defects. To use the standard formulae an idealisation of the defects was made where the geometric shape was simpli®ed to an ellipse. The stress intensity factor of the initial defects,

5. Conclusions 1. It was shown, that the NDT methods worked satisfactorily, 1%CrMoNiV,

Ε = 530 - 550 ˚C

1.5

Mixed mode damage Specimen type (a0/W) D9, D15, C15 (0,2 .. 0,4) Cs25 (0,55) Cs25 (0,55*)

1.0

D30 (0,2)

R″ = ″ n pl / Rmt

Stress - (farfield-) ratio

Ligament damage

D30 (0,4) D60 (0,1) D60 (0,2) D60 (0,4)

Rσ/RK = 2

Cs50 (0,55)

0.5

Crack tip damage

CT100 (0,55)

Rσ/RK = 0,5

Casting with manufact. defects O without crack initiation [24] with crack initiation [24] x large scale round bars [23] 0.0 0.0

0.5

1.0

1.5

Crack tip ratio

2.0

2.5

3.0

3.5

RK = KI id0 / KIi

Fig. 15. Two-criteria-diagram for 1% CrMo(NI)V.steels: initiation point compact tension specimens and DENT-specimens respectively, in comparison with large scale cast steel components with and without crack initiation.

K. Maile et al. / International Journal of Pressure Vessels and Piping 78 (2001) 865±874

Ligament damage

1.0

B C

873

1%CrMoV Cast, E = 530 ˚C

Defect A

Mixed mode damage 4

10 h 3

3

10 h R″ = ″ n pl / Rmt

Stress - (farfield-) ratio

5x10 h

2

10 h 0.5

Crack tip damage

Rσ/RK = 2 Rσ/RK = 0,5

0.0 0.0

0.5

1.0

1.5

2.0

Crack tip ratio RK = KI id0 / KIAC Fig. 16. Two-criteria-diagram for the specimen SFG220 with damage paths for three natural defects.

0.8

creep and creep-fatigue tests = no crack N = below resolution limit crack initiation, lower scatterband below resolution limit / not acceptable

0.6 σ /R n0 p0.2

m

m

0.4

creep-fatigue (cf) creep (c) small sc. specimen, Rm small sc. specimen, Rp1 (cf) crack initiation 2CD (c) crack initiation 2CD

N N

m m

N

crack initiation of the 2CD m = mixed mode damage (other: ligament damage)

0.2 10 1

10 2 10 3 10 time with respect to time at max. load /h

4

Fig. 17. Large scale specimens under creep±fatigue loading, stress related evaluation: normalised net stress versus loading time and ®ndings at internal defects.

both the UT method to indicate internal defects and the MP method to show the opened or partly healed hot tears. 2. In general the results of the SEM investigations show a good accordance with the indication of the C-scan before and at the end of loading. 3. There is a good agreement between the results of the fracture mechanics and the FEM calculation. The

comparison of both analyses shows, that an approximation to describe the defects analytically with an elliptical shape is acceptable. 4. The evaluation with the two-criteria-diagram is conservative. For a ®rst approximation, the Rp1 creep strain curve for defect free specimens can be taken, to estimate the creep- and the creep±fatigue initiation for specimens with defects.

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creep- (c) and creep-fatigue tests (cf)

10

3

m

m

3/2 K (N/mm ) max

m m N

10

10

2

1

10 1

= no crack N = below resolution limit

creep fatigue (cf) creep (c) crack initiation CT25, static (cf) crack initiation 2CD (c) crack initiation 2CD

N

N

crack initiation of the 2CD m = mixed mode damage (other: ligament damage) 3

10 2 10 time with respect to time at max. load /h

10 4

Fig. 18. Large scale specimens under creep±fatigue loading, stress intensity at initial defects versus loading time and ®ndings at defects.

Acknowledgements Thanks are due to the `Bundesminister fuÈr Wirtschaft' (AiF±Nr. 10968) and the `Verein Deutscher Gieûereifachleute' for their support. References [1] Berger C, Mayer K-H, Scarlin RB. In¯uence of natural ¯aws on the operational properties of large forgings for turbogenerators. International Conference on Advances in Material technology for Fossil Power Plants, 1±3 Sept. 1987, Chicago. [2] Berger C, Hollstein T, Mayer K-H, Scarlin RB. Fatigue behaviour of large forgings containing natural defects. Conference on Fatigue 90, 15±20.7.1990, Hawaii. [3] Knipp E. Fehlererscheinungen an GuûstuÈcken. DuÈsseldorf: GiessereiVerlag GmbH, 1953. [4] Colangelo VJ, Heiser FA. Analysis of metallurgical failures. New Nork: Wiley, 1974. [5] Stahl-Eisen-PruÈfblatt 1922. UltraschallpruÈfung von GuûstuÈcken aus ferritischem Stahl, 3. Ausgabe Juli 1985. [6] Stahl-Eisen-PruÈfblatt 1935. Ober¯aÈchenriûpruÈfung von GuûstuÈcken aus Stahl; MagnetpulverpruÈfung, 1. Ausgabe Juni 1982. [7] Euronorm prEN 190/411-1 (draft). [8] Mayer K-H, Berger C, Gnirss G. ZerstoÈrungsfreier Nachweis und

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[11]

[12]

[13] [14] [15] [16] [17]

bruchmechanische Bewertung von herstellungsbedingten Fehlstellen in schweren SchmiedestuÈcken. VDI-Berichte, Heft 917, S. 447/471. Berger C. Ober¯aÈchennahe Fehlstellen im Stahlguû. Ingenieur-Werkstoffe 1991;3(9):55±57. Mayer K-H, Berger C, Gnirss G, Heinrich D, Prestel W. Untersuchungen zur zerstoÈrungsfreien Ermittlung der GroÈûe von natuÈrlichen Fehlstellen in groûen SchmiedestuÈcken von Turbogeneratoren, 17. MPASeminar, Stuttgart, 1991. Mayer K-H, Berger C, Schreiner T, Gerdes C, Goode G. Ultrasonic examination and evaluation of natural defects in large turbomaschinery forging. Fifth International Conference Materials for Advanced Power Engineering, LieÁge, Belgium, October 3±6, 1994. Mayer K-H, Prestel W, Heinrich D. Computer-aided ultrasonic examination and evaluation of natural defects in large turbomachinery forgings for analysis by fracture mechanics. COST 501-Round II, Work package 8, Final Rep 1994. ASME Boiler and Pressure Vessel Code. Section XI. BSI-PD 6493. Guidance on some methods for the derivation of acceptance levels for defects in fusion welded joints, 1980 Mayer KH, Maile K, Gerdes C. Characterisation and quanti®cation of defects in rotors and castings of steam turbines.1st HIDA Conference, Saclay/Paris, France, April 1998 Ewald J, Keienburg K-H. A two-criteria-diagram for creep crack initiation. International Conference on Creep, Tokyo, p.173±8, 14± 18 April 1986. Ewald J, Sheng S, Klenk A, Schellenberg G. Engineering guide to assessment of creep crack initiation on components by two-criteriadiagram. Second HIDA-Conference, Stuttgart, Oct. 2000