- Email: [email protected]
Pressurized Thermal Shock analysis of the reactor pressure vessel
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Procedia Structural Integrity 13 00 (2018) 2083–2088 Structural Integrity Procedia (2016) 000–000
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ECF22 - Loading and Environmental effects on Structural Integrity ECF22 - Loading and Environmental effects on Structural Integrity
Pressurized Thermal Shock analysis of the reactor pressure vessel Pressurized Thermal Shock analysis of the reactor pressure vessel
XV Portuguese Conference on Fracture, PCF 2016, 10-12 February 2016, Paço de Arcos, Portugal
Peter Trampus* Peter Trampus*
Professor emeritus, H-2401 Dunaújváros, University of Dunaújváros, Hungary Thermo-mechanical modeling of a high pressure turbine blade of an Professor emeritus, H-2401 Dunaújváros, University of Dunaújváros, Hungary airplane gas turbine engine Abstract
Abstract a b P.technical Brandão , V.supporting Infantethe , A.M. Deusc* for operation beyond design life. PTS analysis is a substantial part of the analyses license application PTS analysis is aanalysis’ substantial part of the technical analyses supporting the license forthe operation designrules. life. Guidance for the concept is given by the Hungarian nuclear regulator whichapplication is in line with relevantbeyond international a ofsupport Mechanical Engineering, Instituto Superiornuclear Técnico, Universidade Av.with Rovisco Pais, 1,the 1049-001 Lisboa, GuidanceDepartment for the analysis’ concept is given byand the independent Hungarian regulator is in line the relevant international Involving technical organizations consultants as which welldeasLisboa, taking into account changes inrules. fuel Portugal Involving technical support organizations as taking into account changes in fuel management Paks NPP, Hungary, compiled aand newindependent PTS analysis.consultants Subjects ofas thewell analysis were the RPV belt linethe region components b IDMEC, Department of Mechanical Engineering, Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1, 1049-001 Lisboa, management Paks NPP, Hungary, compiled new PTS analysis. Subjects of Beyond the analysis wereinitiating the RPV events belt lineselected region components as well as other circumferential welds of theaRPV including nozzle region. the PTS on the basis Portugal -5 /year. as engineering well as other circumferential welds of the RPV including nozzle region. Beyond the PTS initiating events selected on the Lisboa, basis c of judgment a PSA provided additional transients which showed a higher frequency than 10 Thermal-hydraulic CeFEMA, Department of Mechanical Engineering, Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1, 1049-001 Thermal-hydraulic of engineering judgmentfor a PSA transients showed a higher frequency 10-5 /year. analysis. calculations completed eachprovided selected additional PTS initiating eventwhich provided the necessary input forthan the structural Based on Portugal calculations completed for each selected PTS initiatingNeutron event provided thesurveillance necessary input theused structural analysis. Based on core configurations end-of-life fluence were calculated. dosimetry resultsfor were to verify the calculations. core configurations end-of-life fluence were calculated. Neutron dosimetry surveillance results were used to verify the calculations. Temperature and stress field calculations were performed by solving the system of equations of elasticity. Underclad crack was Temperature stressmechanics field calculations wereKperformed by solving system ofthe equations of between elasticity.the Underclad was/ postulated fracture calculation. at the the crack tip and boundary cladding crack and base I was calculated Abstractforand calculated at the crack tip and the boundary between(Kthe cladding and base / postulated mechanics I was weld metal;for KIcfracture was calculated oncalculation. the basis ofKthe reference curve. Comparison of these two parameters I ≤K Ic), i.e. evaluation weld metal;their KIcpostulated was calculated on the basis of thewas reference curve. Comparison two parameters (KI ≤K Ic), i.e. evaluation of whether the defect stability criteria met, gave theare result. During operation, modern aircraft engine components subjected of to these increasingly demanding operating conditions, of especially whether the defect stability criteria was Such met, gave the result. thepostulated high pressure turbine (HPT) blades. conditions cause these parts to undergo different types of time-dependent © degradation, 2018 The Authors. Published by Elsevier B.V. one of which is AB.V. model using the finite element method (FEM) was developed, in order to be able to predict © 2018 The Published bycreep. Elsevier © 2018creep TheAuthors. Authors. Published by Elsevier B.V. Peer-review under responsibility of the ECF22 organizers. the behaviour of HPT blades. Flight data records (FDR) for a specific aircraft, provided by a commercial aviation Peer-review under responsibility of the ECF22 organizers. Peer-review responsibility the ECF22 company, under were used to obtainofthermal and organizers. mechanical data for three different flight cycles. In order to create the 3D model needed forlifethe FEM analysis, a thermal HPT blade was scanned, and its chemical material properties were Keywords: extension; pressurized shock;scrap plant transients; fast neutron fluence; linearcomposition elastic fractureand mechanics; underclad crack. Keywords: life extension; thermal plant fast neutron fluence; linear elastic fracture mechanics; underclad crack. 3D obtained. The data that pressurized was gathered wasshock; fed into thetransients; FEM model and different simulations were run, first with a simplified rectangular block shape, in order to better establish the model, and then with the real 3D mesh obtained from the blade scrap. The expected behaviour in terms of displacement was observed, in particular at the trailing edge of the blade. Therefore such a 1. overall Introduction can be useful in the goal of predicting turbine blade life, given a set of FDR data. 1. model Introduction
theThe endAuthors. of 2017, the service life extension licensing process of the four VVER-440/V-213 units at Paks NPP, ©At 2016 Published by Elsevier B.V. At the end of 2017, the service life extension licensing of the four VVER-440/V-213 units atperiod Paks NPP, Hungary, was completed. According to the legislation a process formal license application for the extended was Peer-review under responsibility of the Scientific Committee of PCF 2016. Hungary, was completed. According to the legislation a formal license application for the extended period was submitted to the nuclear regulator in which the safe operation had to be demonstrated for the extended term. To justify submitted to the nuclear regulator in which the safe operation had to be demonstrated for the extended term. To justify Highof Pressure Turbine Blade; Creep; Finiteperiod Element(20 Method; 3D Model; Simulation.of the original Time-Limited Ageing theKeywords: feasibility the planned new operation years), the extension the feasibility of the planned new operation period (20 years), the extension of the original Time-Limited Ageing
* Corresponding author. Tel.: +36-20-985-5970 * E-mail Corresponding Tel.: +36-20-985-5970 address:author. [email protected] E-mail address: [email protected] 2452-3216 © 2018 The Authors. Published by Elsevier B.V. 2452-3216 © 2018 Authors. Published Elsevier B.V. Peer-review underThe responsibility of theby ECF22 organizers. * Corresponding Tel.: +351of218419991. Peer-review underauthor. responsibility the ECF22 organizers. E-mail address: [email protected] 2452-3216 © 2016 The Authors. Published by Elsevier B.V.
Peer-review under responsibility of the Scientific Committee of PCF 2016. 2452-3216 2018 The Authors. Published by Elsevier B.V. Peer-review under responsibility of the ECF22 organizers. 10.1016/j.prostr.2018.12.204
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Peter Trampus / Procedia Structural Integrity 13 (2018) 2083–2088 P. Trampus / Structural Integrity Procedia 00 (2018) 000–000
Analyses (TLAA) valid for 30 years was also requested by the regulator. Among TLAA the most significant analysis relates to the Pressurized Thermal Shock (PTS) for the reactor pressure vessels (RPVs). The RPV is the only nonreplaceable component determining thus the plant’s technically feasible lifespan. PTS is an overcooling transient which causes a thermal shock to the RPV, while the pressure is either maintained or the system is re-pressurized during the transient. The thermal stress due to the rapid cooling of the vessel wall in combination with the pressure stress results in large tensile stresses which have their maximum value in the inside surface of the vessel. Also, a unique ageing phenomenon called irradiation embrittlement occurs in the RPV wall reducing the structural material’s fracture toughness and shifting the ductile-brittle transition temperature in the direction of higher temperature. In the case if a flaw (crack) would exist in an area of the vessel wall near to the inside surface where the material properties degraded due to fast neutron irradiation and a PTS transient would happen, the RPV integrity would be jeopardized, see the process in Fig. 1.
Fig. 1. PTS process and its impact.
The goal of this paper is to describe the methodology and the basic results of the PTS analysis performed for Paks NPP. It gives an example for the effect of loading and environment on the structural integrity of a large-scale pressurized component. Leading technical support organizations and both Hungarian and international consultants contributed to the accomplishment and independent validation of the analysis. The author chaired the expert committee dealing with RPV integrity as well as contributed to and edited the report on PTS analysis submitted to the regulator. 2. Inputs for PTS calculation The PTS analysis followed the Hungarian regulation which is in line with the guidance of the International Atomic Energy Agency (IAEA 2006). The structural integrity against brittle fracture of the RPV is ensured if the factual ductile-brittle transition temperature, called critical temperature of brittleness Tk , of its critical components is less allow . The analysis is based on the than the maximum allowable component specific transition temperature Tk comparison of the static fracture toughness K Ic of the structural material and the stress intensity factor K I calculated from the given loading situation using the theory of linear-elastic fracture mechanics. Subjects of the PTS analysis were all possible RPV locations that could potentially be susceptible for service induced ageing. However, the focus of the analysis was placed on the RPV belt line region (base metal of forged ring and circumferential core weld) because of the intensive embrittlement process caused by fast neutrons. The major steps of the overall methodology of PTS calculations are as follows: • identification of the PTS transients, • thermal-hydraulic analyses of PTS transients, • neutron fluence calculations, • evaluation of material properties and irradiation effects on RPV structural materials,
Peter Trampus / Procedia Structural Integrity 13 (2018) 2083–2088 P. Trampus / Structural Integrity Procedia 00 (2018) 000–000
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• temperature and strength calculations, • fracture mechanics evaluation, • determination of allowed service life for RPVs. 2.1. PTS transient identification The identification of PTS transients has been performed in a comprehensive way taking into account various accident sequences including impact of component malfunctions, different operator actions, internal and external hazards. Eleven specified transients were chosen and analyzed on deterministic basis, i.e. engineering judgment. Complementary to this, selection of additional transients on probabilistic basis was also performed. As for the latter those events were selected which showed a higher probability of occurrence than 10-5/year. 2.2. Thermal-hydraulic analyses of PTS transients The overall progression of accidents was calculated with an advanced thermal-hydraulic system codes (RELAP5/mod3.2 and ATHLET). The adequate margin of the calculation results was ensured by the use of bestestimate computer codes with conservative input data and with conservative assumptions regarding the availabilities of influencing systems. In flow stagnation cases the role of the system code calculations was to estimate the onset of the stagnation, and give the initial temperature conditions, emergency core cooling injection flow rates and the so called well-mixed temperatures in the RPV down-comer segments while the actual temperature curves for the so called colder plumes in these cases were calculated with a separate code. 2.3. Neutron fluence calculations Based on core configurations implemented until current analysis and planned to be implemented in the future, calculations using the core design code and Monte Carlo transport code were performed, and end-of-life fluence FEOL for 50 operating years, were calculated for the RPV wall as well as for the surveillance position. Neutron dosimetry surveillance results were used to verify the calculations. The calculations were made in the azimuthal and radial position where the fluence has its maximal value. Table 1 shows the results of calculations for the end-of-life fluence for RPVs 1 to 4 for the targeted operating lifetime (50 years). Table 1. End-of-life fast neutron fluence [n/cm2] (E>0.5MeV) RPV
Core weld
Core center
1
1.94E+20
2.81E+20
2
1.97E+20
2.83E+20
3
1.89E+20
2.72E+20
4
1.95E+20
2.79E+20
3. Irradiation effects on RPV materials The RPV shells were made of forgings from Cr-Mo-V alloyed steel 15Ch2MFA. The welding of separate shells to one another was carried out with the use of submerged method with Sv-10ChMFT wire and the flux AN-42. The anticorrosive cladding of RPV inner surface has two layers: the first layer is Sv-07Cr25Ni13; the second layer, first bead is Sv-08Cr19Ni10Mn2Nb and second bead is Sv-04Cr20Ni10MnNb. The fracture toughness reference curve of RPV materials 15Ch2MFA and Sv-10ChMFT is as follows: 𝐾𝐾𝐼𝐼𝐼𝐼 = 𝑚𝑚𝑚𝑚𝑚𝑚{26 + 36𝑒𝑒𝑒𝑒𝑒𝑒[0.02(𝑇𝑇 − 𝑇𝑇𝑘𝑘 )], 200},
in MPa√m,
(1)
where: T is the material’s temperature and Tk is the critical temperature of brittleness. During evaluation of the shift of ΔTk the effects of irradiation, thermal ageing and fatigue were taken into account. Tk and ΔTk were determined during evaluation of Charpy impact test results of surveillance specimens for every RPV. As for the initial value of the critical temperature of brittleness Tk0 the more conservative value of measurements performed by the manufacturer and performed during surveillance programs has been considered. The zero level measurements of surveillance programs are evaluated using the 41 J impact energy criterion, similarly to the evaluation of tests performed on irradiated specimens.
Peter Trampus / Procedia Structural Integrity 13 (2018) 2083–2088 P. Trampus / Structural Integrity Procedia 00 (2018) 000–000
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The transition temperature shifts due to fast neutron irradiation (ΔTF) were determined by tangent hyperbolic curve fitting to Charpy impact test results of the given set applying the 41 J criterion. The transition temperature shift curves depending on the neutron fluence were derived from a two-parameter fitting as follows: ∆𝑇𝑇𝐹𝐹 (𝐹𝐹𝑛𝑛 ) = 𝐴𝐴𝐹𝐹 {𝐹𝐹𝑛𝑛 ⁄𝐹𝐹0 }𝑛𝑛 ,
in °C,
TT = Tkt − Tk 0 ,
in °C,
(2)
where: Fn is the fast neutron fluence (E>0.5 MeV) calculated for the surveillance specimens by the neutron fluence calculations, F=1020 n/cm2; AF and n are the parameters searched during fitting. Based on experimental temperature measures, e.g. (Ballesteros et al 2003), specimen overheating in capsules of irradiation chains was not considered. During research reported in (IAEA 2010), it was established that in the case if copper content of the structural material is low, and the lead factor is greater than that recommended in relevant standard (ASTM E 185) and the irradiation temperature does not exceed 290 °C, flux effect should not be considered. The copper content of the RPV materials at Paks NPP is low, the lead factor is much higher than the range recommended in ASTM E 185 (~12 vs 1 to 3) and the irradiation temperature is not higher than 270 °C; however, based upon current understanding and limited data, the effect of neutron flux cannot be determined and is assumed to be negligible. The transition temperature shifts of the RPV critical elements due to thermal embrittlement ΔTT were determined on the basis of the results of analyses performed on unit specific surveillance specimens where thermally aged specimens were included into the chains. The transition temperature shift for thermal embrittlement is as follows: (3)
where: Tkt is the transition temperature of the material after thermal embrittlement. Sensitivity to thermal ageing of materials with low copper content such as 15Ch2MFA and Sv-10ChMFT is negligible in the range of operating temperature (~270 °C). The transition temperature shifts of the RPV critical elements due to fatigue ΔTN could be neglected for the belt line region of the RPVs. This is because there are no stress concentrations here, and the numbers and slope of heating up and cooling down cycles are limited. The critical temperature of brittleness Tk of the RPV critical elements was determined according the following formula: 𝑻𝑻𝒌𝒌 =𝑻𝑻𝒌𝒌𝒌𝒌 + ∆𝑻𝑻𝑭𝑭 + ∆𝑻𝑻𝑻𝑻 + ∆𝑻𝑻𝑵𝑵 + 𝝈𝝈,
in °C,
(4)
The uncertainty range of 1∙σ (10 °C for forgings, and 16 °C for welds). As an example, the equation of the Tk (Fn ) trend curve developed for RPV 1 weld metal is as follows: Tk = 25.0 + 62.364(Fn/1020)0.392+16,
in °C.
(5)
The Tk values of RPV 1 weld metal (core weld) as a function of fast neutron fluence, i.e. the trend curve, pertaining to service life up to the 50th year is shown in Fig 2. 4. Evaluation of RPV in-service inspection results As part of the in-service inspection (ISI) program, standardized non-destructive testing (NDT) methods such as visual testing (VT), ultrasonic testing (UT), eddy current testing (ET), liquid penetration testing (PT) and acousticemission testing (AET) have been applied to the parts of the RPV that are important from the point of view of integrity or nuclear safety, and performed manually or by remotely controlled manipulator according to the accessibility and radiation conditions of the given location. UT is implemented from both the inside diameter (ID) and outside diameter (OD) of the RPV with 10 years periodicity. ID inspection has always been done by vendor whilst examination from OD is performed by the plant staff using the upgraded USK-213 equipment (originally Russian design). The ID UT system (equipment, procedure) applied for examination of circumferential welds, base metal in beltline region, nozzle inner radii and safe end-to-pipe dissimilar welds, and the ET system (equipment, procedure) applied for examination of cladding are qualified in accordance with the European methodology (ENIQ 2007). The ISI results showed that the integrity of the RPV with regard to the areas within the scope of ISI NDT of the circumferential welds, the belt line material and the cladding is proven
Peter Trampus / Procedia Structural Integrity 13 (2018) 2083–2088 P. Trampus / Structural Integrity Procedia 00 (2018) 000–000
2087 5
Tk trend curves for weld metal (No. 5/6) of Unit 1 based on surveillance data 180 Fl. Calculated Fl. Measured Fitted Tk
160
o
Fitted Tk +16 C
140
Tk [°C]
120 100 Data: CV1A_B Model: user31
80
Chi^2/DoF = 9.11428 R^2 = 0.98323
60
a b Tk0
40
62.36351 0.39205 o 25 C
±2.59772 ±0.03846
20 0
1
2
3 20
4
5
2
Fluence [*10 n/cm E>0.5 MeV]
.
Fig. 2. Trend curve of RPV 1 weld metal.
5. Fracture mechanics analysis All the transients, selected by PTS event screening criterion were the subject of fracture mechanics analysis. Taking the loads (pressure, temperature and heat transfer coefficient) resulting from thermal-hydraulic analyses for the particular PTS events into account as boundary conditions, the global temperature, strain and stress fields were firstly evaluated using 3D finite element (FE) models of the RPV without crack, assuming linear elastic material behavior. After having the stress field, a set of underclad cracks of semi-elliptic type were postulated for calculation of the stress intensity factors KI with the maximum depth of a = 0.1t and with an aspect ratio of a/c = 1/3. In the base metal the postulated crack orientation was normal to the principal stress and in the circumferential weld it was circumferential. The stress intensity factor values were calculated at the deepest point of the crack and at the near interface location of the cladding and the base / weld metal using an engineering evaluation method (Marie and Chapuliot 2008). Tkallow values for selected locations of flaws were then evaluated. KIc was computed using Tk and the calculated through wall temperature distribution for each flaw depth and evaluation point on the flaw boundary as a function of time during the transient. Then, KI and KIc was plotted as a function of temperature for each flaw depth and evaluation point on the flaw boundary for each time during the transient. Each KIc vs temperature curve was moved along the KI axis until it was just tangent to the associated curve. The Tk value at which the KIc vs temperature curve was just tangent to the associated KI vs temperature curve was Tkallow for the selected flaw depth and evaluation point. Tkallow value for a designed location was defined as the minimum of Tkallow values computed for all flaw depths and all evaluation points defined for that location. For demonstration of the results, one of the most severe transients is shown here: Ø492 mm hot leg line, double ended guillotine brake (large break loss-of-coolant accident, LB LOCA). Fig. 3 shows KI (as a function of local temperature) at the deepest point of the crack front and at the closest point to the cladding-base metal interface. The same figure shows the KIc reference curve belonging to Tkallow value determined for loads appearing at these points.
Peter Trampus / Procedia Structural Integrity 13 (2018) 2083–2088 P. Trampus / Structural Integrity Procedia 00 (2018) 000–000
2088 6 (MPam)
KI Point 1 KI Point 6 KIc Point 1
K I values for the deepest (1) and near interface point (6) of a 23 mm deep crack (a/c=14/42 mm) vs K IC tangent curve
100 90 80 70 60 50 40 30 20 10 0 0
20
40
60
80
100
120
140
160
180
200
220
240
260
280
(°C)
Fig. 3. KI and KIc values: deepest point (1) and point nearest to the cladding-base metal boundary (6), during the transient LB LOCA.
Tkallow value for a component was defined as the minimum of Tkallow values calculated in various circumferential sectors of the RPVs. The allowable service life (τallow) of a component is the maximum lifetime τ – measured in service years – for which Tk(τ) ≤ Tkallow. The allowable service life of an RPV is the minimum of τallow values defined for components selected for fracture mechanics analyses. The results based on fracture mechanics analyses admit more than 60 years of operation for the allowable lifetime of RPVs 1 to 4 at Paks NPP. Acknowledgements The author would like to thank Messrs. G. Bona, J. Elter and S. Ratkai from Paks NPP, T. Fekete, F. Gillemot and A. Kereszturi from Centre for Energy Research, Hungarian Academy of Sciences, and Ms. E. M. Zsolnay from Institute for Nuclear Techniques, Budapest University of Technology and Economics, for their invaluable contribution to this work. References ASTM E185-16, 2016. Standard Practice for Design of Programs for Light-Water Moderated Nuclear Power Reactor Vessels, ASTM International, West Conshohocken, PA Ballesteros, A. et al, 2003. Assessment of irradiation conditions in WWER-440 (213) RPV surveillance location (COBRA Project), In: Proc. Int. Seminar Networking for Effective R&D, Petten, Report EUR 20984 EN ENIQ Guideline, 2007. European Methodology for Qualification of Non-Destructive Testing (third issue), EUR 17299 EN, Luxembourg IAEA-EBP-WWER-08, 2006. Guidelines on Pressurized Thermal Shock Analysis for WWER Nuclear Power Plants (Rev. 1), IAEA, Vienna IAEA TECDOC 1627, 2010. Pressurized Thermal Shock in Nuclear Power Plants: Good Practices for Assessment, , IAEA, Vienna Marie, S., Chapuliot, S., 2008. Improvement of the calculation of the stress intensity factor for underclad and trough-clad defects in a reactor pressure vessel subjected to pressurized thermal shock. Int. J. Pressure Vessels and Piping 85, 517-531.
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Structural Integrity Procedia 00 (2018) 000–000 Available online www.sciencedirect.com Available online at at www.sciencedirect.com Structural Integrity Procedia 00 (2018) 000–000
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Procedia Structural Integrity 13 00 (2018) 2083–2088 Structural Integrity Procedia (2016) 000–000
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ECF22 - Loading and Environmental effects on Structural Integrity ECF22 - Loading and Environmental effects on Structural Integrity
Pressurized Thermal Shock analysis of the reactor pressure vessel Pressurized Thermal Shock analysis of the reactor pressure vessel
XV Portuguese Conference on Fracture, PCF 2016, 10-12 February 2016, Paço de Arcos, Portugal
Peter Trampus* Peter Trampus*
Professor emeritus, H-2401 Dunaújváros, University of Dunaújváros, Hungary Thermo-mechanical modeling of a high pressure turbine blade of an Professor emeritus, H-2401 Dunaújváros, University of Dunaújváros, Hungary airplane gas turbine engine Abstract
Abstract a b P.technical Brandão , V.supporting Infantethe , A.M. Deusc* for operation beyond design life. PTS analysis is a substantial part of the analyses license application PTS analysis is aanalysis’ substantial part of the technical analyses supporting the license forthe operation designrules. life. Guidance for the concept is given by the Hungarian nuclear regulator whichapplication is in line with relevantbeyond international a ofsupport Mechanical Engineering, Instituto Superiornuclear Técnico, Universidade Av.with Rovisco Pais, 1,the 1049-001 Lisboa, GuidanceDepartment for the analysis’ concept is given byand the independent Hungarian regulator is in line the relevant international Involving technical organizations consultants as which welldeasLisboa, taking into account changes inrules. fuel Portugal Involving technical support organizations as taking into account changes in fuel management Paks NPP, Hungary, compiled aand newindependent PTS analysis.consultants Subjects ofas thewell analysis were the RPV belt linethe region components b IDMEC, Department of Mechanical Engineering, Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1, 1049-001 Lisboa, management Paks NPP, Hungary, compiled new PTS analysis. Subjects of Beyond the analysis wereinitiating the RPV events belt lineselected region components as well as other circumferential welds of theaRPV including nozzle region. the PTS on the basis Portugal -5 /year. as engineering well as other circumferential welds of the RPV including nozzle region. Beyond the PTS initiating events selected on the Lisboa, basis c of judgment a PSA provided additional transients which showed a higher frequency than 10 Thermal-hydraulic CeFEMA, Department of Mechanical Engineering, Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1, 1049-001 Thermal-hydraulic of engineering judgmentfor a PSA transients showed a higher frequency 10-5 /year. analysis. calculations completed eachprovided selected additional PTS initiating eventwhich provided the necessary input forthan the structural Based on Portugal calculations completed for each selected PTS initiatingNeutron event provided thesurveillance necessary input theused structural analysis. Based on core configurations end-of-life fluence were calculated. dosimetry resultsfor were to verify the calculations. core configurations end-of-life fluence were calculated. Neutron dosimetry surveillance results were used to verify the calculations. Temperature and stress field calculations were performed by solving the system of equations of elasticity. Underclad crack was Temperature stressmechanics field calculations wereKperformed by solving system ofthe equations of between elasticity.the Underclad was/ postulated fracture calculation. at the the crack tip and boundary cladding crack and base I was calculated Abstractforand calculated at the crack tip and the boundary between(Kthe cladding and base / postulated mechanics I was weld metal;for KIcfracture was calculated oncalculation. the basis ofKthe reference curve. Comparison of these two parameters I ≤K Ic), i.e. evaluation weld metal;their KIcpostulated was calculated on the basis of thewas reference curve. Comparison two parameters (KI ≤K Ic), i.e. evaluation of whether the defect stability criteria met, gave theare result. During operation, modern aircraft engine components subjected of to these increasingly demanding operating conditions, of especially whether the defect stability criteria was Such met, gave the result. thepostulated high pressure turbine (HPT) blades. conditions cause these parts to undergo different types of time-dependent © degradation, 2018 The Authors. Published by Elsevier B.V. one of which is AB.V. model using the finite element method (FEM) was developed, in order to be able to predict © 2018 The Published bycreep. Elsevier © 2018creep TheAuthors. Authors. Published by Elsevier B.V. Peer-review under responsibility of the ECF22 organizers. the behaviour of HPT blades. Flight data records (FDR) for a specific aircraft, provided by a commercial aviation Peer-review under responsibility of the ECF22 organizers. Peer-review responsibility the ECF22 company, under were used to obtainofthermal and organizers. mechanical data for three different flight cycles. In order to create the 3D model needed forlifethe FEM analysis, a thermal HPT blade was scanned, and its chemical material properties were Keywords: extension; pressurized shock;scrap plant transients; fast neutron fluence; linearcomposition elastic fractureand mechanics; underclad crack. Keywords: life extension; thermal plant fast neutron fluence; linear elastic fracture mechanics; underclad crack. 3D obtained. The data that pressurized was gathered wasshock; fed into thetransients; FEM model and different simulations were run, first with a simplified rectangular block shape, in order to better establish the model, and then with the real 3D mesh obtained from the blade scrap. The expected behaviour in terms of displacement was observed, in particular at the trailing edge of the blade. Therefore such a 1. overall Introduction can be useful in the goal of predicting turbine blade life, given a set of FDR data. 1. model Introduction
theThe endAuthors. of 2017, the service life extension licensing process of the four VVER-440/V-213 units at Paks NPP, ©At 2016 Published by Elsevier B.V. At the end of 2017, the service life extension licensing of the four VVER-440/V-213 units atperiod Paks NPP, Hungary, was completed. According to the legislation a process formal license application for the extended was Peer-review under responsibility of the Scientific Committee of PCF 2016. Hungary, was completed. According to the legislation a formal license application for the extended period was submitted to the nuclear regulator in which the safe operation had to be demonstrated for the extended term. To justify submitted to the nuclear regulator in which the safe operation had to be demonstrated for the extended term. To justify Highof Pressure Turbine Blade; Creep; Finiteperiod Element(20 Method; 3D Model; Simulation.of the original Time-Limited Ageing theKeywords: feasibility the planned new operation years), the extension the feasibility of the planned new operation period (20 years), the extension of the original Time-Limited Ageing
* Corresponding author. Tel.: +36-20-985-5970 * E-mail Corresponding Tel.: +36-20-985-5970 address:author. [email protected] E-mail address: [email protected] 2452-3216 © 2018 The Authors. Published by Elsevier B.V. 2452-3216 © 2018 Authors. Published Elsevier B.V. Peer-review underThe responsibility of theby ECF22 organizers. * Corresponding Tel.: +351of218419991. Peer-review underauthor. responsibility the ECF22 organizers. E-mail address: [email protected] 2452-3216 © 2016 The Authors. Published by Elsevier B.V.
Peer-review under responsibility of the Scientific Committee of PCF 2016. 2452-3216 2018 The Authors. Published by Elsevier B.V. Peer-review under responsibility of the ECF22 organizers. 10.1016/j.prostr.2018.12.204
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Peter Trampus / Procedia Structural Integrity 13 (2018) 2083–2088 P. Trampus / Structural Integrity Procedia 00 (2018) 000–000
Analyses (TLAA) valid for 30 years was also requested by the regulator. Among TLAA the most significant analysis relates to the Pressurized Thermal Shock (PTS) for the reactor pressure vessels (RPVs). The RPV is the only nonreplaceable component determining thus the plant’s technically feasible lifespan. PTS is an overcooling transient which causes a thermal shock to the RPV, while the pressure is either maintained or the system is re-pressurized during the transient. The thermal stress due to the rapid cooling of the vessel wall in combination with the pressure stress results in large tensile stresses which have their maximum value in the inside surface of the vessel. Also, a unique ageing phenomenon called irradiation embrittlement occurs in the RPV wall reducing the structural material’s fracture toughness and shifting the ductile-brittle transition temperature in the direction of higher temperature. In the case if a flaw (crack) would exist in an area of the vessel wall near to the inside surface where the material properties degraded due to fast neutron irradiation and a PTS transient would happen, the RPV integrity would be jeopardized, see the process in Fig. 1.
Fig. 1. PTS process and its impact.
The goal of this paper is to describe the methodology and the basic results of the PTS analysis performed for Paks NPP. It gives an example for the effect of loading and environment on the structural integrity of a large-scale pressurized component. Leading technical support organizations and both Hungarian and international consultants contributed to the accomplishment and independent validation of the analysis. The author chaired the expert committee dealing with RPV integrity as well as contributed to and edited the report on PTS analysis submitted to the regulator. 2. Inputs for PTS calculation The PTS analysis followed the Hungarian regulation which is in line with the guidance of the International Atomic Energy Agency (IAEA 2006). The structural integrity against brittle fracture of the RPV is ensured if the factual ductile-brittle transition temperature, called critical temperature of brittleness Tk , of its critical components is less allow . The analysis is based on the than the maximum allowable component specific transition temperature Tk comparison of the static fracture toughness K Ic of the structural material and the stress intensity factor K I calculated from the given loading situation using the theory of linear-elastic fracture mechanics. Subjects of the PTS analysis were all possible RPV locations that could potentially be susceptible for service induced ageing. However, the focus of the analysis was placed on the RPV belt line region (base metal of forged ring and circumferential core weld) because of the intensive embrittlement process caused by fast neutrons. The major steps of the overall methodology of PTS calculations are as follows: • identification of the PTS transients, • thermal-hydraulic analyses of PTS transients, • neutron fluence calculations, • evaluation of material properties and irradiation effects on RPV structural materials,
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• temperature and strength calculations, • fracture mechanics evaluation, • determination of allowed service life for RPVs. 2.1. PTS transient identification The identification of PTS transients has been performed in a comprehensive way taking into account various accident sequences including impact of component malfunctions, different operator actions, internal and external hazards. Eleven specified transients were chosen and analyzed on deterministic basis, i.e. engineering judgment. Complementary to this, selection of additional transients on probabilistic basis was also performed. As for the latter those events were selected which showed a higher probability of occurrence than 10-5/year. 2.2. Thermal-hydraulic analyses of PTS transients The overall progression of accidents was calculated with an advanced thermal-hydraulic system codes (RELAP5/mod3.2 and ATHLET). The adequate margin of the calculation results was ensured by the use of bestestimate computer codes with conservative input data and with conservative assumptions regarding the availabilities of influencing systems. In flow stagnation cases the role of the system code calculations was to estimate the onset of the stagnation, and give the initial temperature conditions, emergency core cooling injection flow rates and the so called well-mixed temperatures in the RPV down-comer segments while the actual temperature curves for the so called colder plumes in these cases were calculated with a separate code. 2.3. Neutron fluence calculations Based on core configurations implemented until current analysis and planned to be implemented in the future, calculations using the core design code and Monte Carlo transport code were performed, and end-of-life fluence FEOL for 50 operating years, were calculated for the RPV wall as well as for the surveillance position. Neutron dosimetry surveillance results were used to verify the calculations. The calculations were made in the azimuthal and radial position where the fluence has its maximal value. Table 1 shows the results of calculations for the end-of-life fluence for RPVs 1 to 4 for the targeted operating lifetime (50 years). Table 1. End-of-life fast neutron fluence [n/cm2] (E>0.5MeV) RPV
Core weld
Core center
1
1.94E+20
2.81E+20
2
1.97E+20
2.83E+20
3
1.89E+20
2.72E+20
4
1.95E+20
2.79E+20
3. Irradiation effects on RPV materials The RPV shells were made of forgings from Cr-Mo-V alloyed steel 15Ch2MFA. The welding of separate shells to one another was carried out with the use of submerged method with Sv-10ChMFT wire and the flux AN-42. The anticorrosive cladding of RPV inner surface has two layers: the first layer is Sv-07Cr25Ni13; the second layer, first bead is Sv-08Cr19Ni10Mn2Nb and second bead is Sv-04Cr20Ni10MnNb. The fracture toughness reference curve of RPV materials 15Ch2MFA and Sv-10ChMFT is as follows: 𝐾𝐾𝐼𝐼𝐼𝐼 = 𝑚𝑚𝑚𝑚𝑚𝑚{26 + 36𝑒𝑒𝑒𝑒𝑒𝑒[0.02(𝑇𝑇 − 𝑇𝑇𝑘𝑘 )], 200},
in MPa√m,
(1)
where: T is the material’s temperature and Tk is the critical temperature of brittleness. During evaluation of the shift of ΔTk the effects of irradiation, thermal ageing and fatigue were taken into account. Tk and ΔTk were determined during evaluation of Charpy impact test results of surveillance specimens for every RPV. As for the initial value of the critical temperature of brittleness Tk0 the more conservative value of measurements performed by the manufacturer and performed during surveillance programs has been considered. The zero level measurements of surveillance programs are evaluated using the 41 J impact energy criterion, similarly to the evaluation of tests performed on irradiated specimens.
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The transition temperature shifts due to fast neutron irradiation (ΔTF) were determined by tangent hyperbolic curve fitting to Charpy impact test results of the given set applying the 41 J criterion. The transition temperature shift curves depending on the neutron fluence were derived from a two-parameter fitting as follows: ∆𝑇𝑇𝐹𝐹 (𝐹𝐹𝑛𝑛 ) = 𝐴𝐴𝐹𝐹 {𝐹𝐹𝑛𝑛 ⁄𝐹𝐹0 }𝑛𝑛 ,
in °C,
TT = Tkt − Tk 0 ,
in °C,
(2)
where: Fn is the fast neutron fluence (E>0.5 MeV) calculated for the surveillance specimens by the neutron fluence calculations, F=1020 n/cm2; AF and n are the parameters searched during fitting. Based on experimental temperature measures, e.g. (Ballesteros et al 2003), specimen overheating in capsules of irradiation chains was not considered. During research reported in (IAEA 2010), it was established that in the case if copper content of the structural material is low, and the lead factor is greater than that recommended in relevant standard (ASTM E 185) and the irradiation temperature does not exceed 290 °C, flux effect should not be considered. The copper content of the RPV materials at Paks NPP is low, the lead factor is much higher than the range recommended in ASTM E 185 (~12 vs 1 to 3) and the irradiation temperature is not higher than 270 °C; however, based upon current understanding and limited data, the effect of neutron flux cannot be determined and is assumed to be negligible. The transition temperature shifts of the RPV critical elements due to thermal embrittlement ΔTT were determined on the basis of the results of analyses performed on unit specific surveillance specimens where thermally aged specimens were included into the chains. The transition temperature shift for thermal embrittlement is as follows: (3)
where: Tkt is the transition temperature of the material after thermal embrittlement. Sensitivity to thermal ageing of materials with low copper content such as 15Ch2MFA and Sv-10ChMFT is negligible in the range of operating temperature (~270 °C). The transition temperature shifts of the RPV critical elements due to fatigue ΔTN could be neglected for the belt line region of the RPVs. This is because there are no stress concentrations here, and the numbers and slope of heating up and cooling down cycles are limited. The critical temperature of brittleness Tk of the RPV critical elements was determined according the following formula: 𝑻𝑻𝒌𝒌 =𝑻𝑻𝒌𝒌𝒌𝒌 + ∆𝑻𝑻𝑭𝑭 + ∆𝑻𝑻𝑻𝑻 + ∆𝑻𝑻𝑵𝑵 + 𝝈𝝈,
in °C,
(4)
The uncertainty range of 1∙σ (10 °C for forgings, and 16 °C for welds). As an example, the equation of the Tk (Fn ) trend curve developed for RPV 1 weld metal is as follows: Tk = 25.0 + 62.364(Fn/1020)0.392+16,
in °C.
(5)
The Tk values of RPV 1 weld metal (core weld) as a function of fast neutron fluence, i.e. the trend curve, pertaining to service life up to the 50th year is shown in Fig 2. 4. Evaluation of RPV in-service inspection results As part of the in-service inspection (ISI) program, standardized non-destructive testing (NDT) methods such as visual testing (VT), ultrasonic testing (UT), eddy current testing (ET), liquid penetration testing (PT) and acousticemission testing (AET) have been applied to the parts of the RPV that are important from the point of view of integrity or nuclear safety, and performed manually or by remotely controlled manipulator according to the accessibility and radiation conditions of the given location. UT is implemented from both the inside diameter (ID) and outside diameter (OD) of the RPV with 10 years periodicity. ID inspection has always been done by vendor whilst examination from OD is performed by the plant staff using the upgraded USK-213 equipment (originally Russian design). The ID UT system (equipment, procedure) applied for examination of circumferential welds, base metal in beltline region, nozzle inner radii and safe end-to-pipe dissimilar welds, and the ET system (equipment, procedure) applied for examination of cladding are qualified in accordance with the European methodology (ENIQ 2007). The ISI results showed that the integrity of the RPV with regard to the areas within the scope of ISI NDT of the circumferential welds, the belt line material and the cladding is proven
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Tk trend curves for weld metal (No. 5/6) of Unit 1 based on surveillance data 180 Fl. Calculated Fl. Measured Fitted Tk
160
o
Fitted Tk +16 C
140
Tk [°C]
120 100 Data: CV1A_B Model: user31
80
Chi^2/DoF = 9.11428 R^2 = 0.98323
60
a b Tk0
40
62.36351 0.39205 o 25 C
±2.59772 ±0.03846
20 0
1
2
3 20
4
5
2
Fluence [*10 n/cm E>0.5 MeV]
.
Fig. 2. Trend curve of RPV 1 weld metal.
5. Fracture mechanics analysis All the transients, selected by PTS event screening criterion were the subject of fracture mechanics analysis. Taking the loads (pressure, temperature and heat transfer coefficient) resulting from thermal-hydraulic analyses for the particular PTS events into account as boundary conditions, the global temperature, strain and stress fields were firstly evaluated using 3D finite element (FE) models of the RPV without crack, assuming linear elastic material behavior. After having the stress field, a set of underclad cracks of semi-elliptic type were postulated for calculation of the stress intensity factors KI with the maximum depth of a = 0.1t and with an aspect ratio of a/c = 1/3. In the base metal the postulated crack orientation was normal to the principal stress and in the circumferential weld it was circumferential. The stress intensity factor values were calculated at the deepest point of the crack and at the near interface location of the cladding and the base / weld metal using an engineering evaluation method (Marie and Chapuliot 2008). Tkallow values for selected locations of flaws were then evaluated. KIc was computed using Tk and the calculated through wall temperature distribution for each flaw depth and evaluation point on the flaw boundary as a function of time during the transient. Then, KI and KIc was plotted as a function of temperature for each flaw depth and evaluation point on the flaw boundary for each time during the transient. Each KIc vs temperature curve was moved along the KI axis until it was just tangent to the associated curve. The Tk value at which the KIc vs temperature curve was just tangent to the associated KI vs temperature curve was Tkallow for the selected flaw depth and evaluation point. Tkallow value for a designed location was defined as the minimum of Tkallow values computed for all flaw depths and all evaluation points defined for that location. For demonstration of the results, one of the most severe transients is shown here: Ø492 mm hot leg line, double ended guillotine brake (large break loss-of-coolant accident, LB LOCA). Fig. 3 shows KI (as a function of local temperature) at the deepest point of the crack front and at the closest point to the cladding-base metal interface. The same figure shows the KIc reference curve belonging to Tkallow value determined for loads appearing at these points.
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KI Point 1 KI Point 6 KIc Point 1
K I values for the deepest (1) and near interface point (6) of a 23 mm deep crack (a/c=14/42 mm) vs K IC tangent curve
100 90 80 70 60 50 40 30 20 10 0 0
20
40
60
80
100
120
140
160
180
200
220
240
260
280
(°C)
Fig. 3. KI and KIc values: deepest point (1) and point nearest to the cladding-base metal boundary (6), during the transient LB LOCA.
Tkallow value for a component was defined as the minimum of Tkallow values calculated in various circumferential sectors of the RPVs. The allowable service life (τallow) of a component is the maximum lifetime τ – measured in service years – for which Tk(τ) ≤ Tkallow. The allowable service life of an RPV is the minimum of τallow values defined for components selected for fracture mechanics analyses. The results based on fracture mechanics analyses admit more than 60 years of operation for the allowable lifetime of RPVs 1 to 4 at Paks NPP. Acknowledgements The author would like to thank Messrs. G. Bona, J. Elter and S. Ratkai from Paks NPP, T. Fekete, F. Gillemot and A. Kereszturi from Centre for Energy Research, Hungarian Academy of Sciences, and Ms. E. M. Zsolnay from Institute for Nuclear Techniques, Budapest University of Technology and Economics, for their invaluable contribution to this work. References ASTM E185-16, 2016. Standard Practice for Design of Programs for Light-Water Moderated Nuclear Power Reactor Vessels, ASTM International, West Conshohocken, PA Ballesteros, A. et al, 2003. Assessment of irradiation conditions in WWER-440 (213) RPV surveillance location (COBRA Project), In: Proc. Int. Seminar Networking for Effective R&D, Petten, Report EUR 20984 EN ENIQ Guideline, 2007. European Methodology for Qualification of Non-Destructive Testing (third issue), EUR 17299 EN, Luxembourg IAEA-EBP-WWER-08, 2006. Guidelines on Pressurized Thermal Shock Analysis for WWER Nuclear Power Plants (Rev. 1), IAEA, Vienna IAEA TECDOC 1627, 2010. Pressurized Thermal Shock in Nuclear Power Plants: Good Practices for Assessment, , IAEA, Vienna Marie, S., Chapuliot, S., 2008. Improvement of the calculation of the stress intensity factor for underclad and trough-clad defects in a reactor pressure vessel subjected to pressurized thermal shock. Int. J. Pressure Vessels and Piping 85, 517-531.