Pre-qualification of additively manufactured 316L stainless steel using small punch and uniaxial creep testing

Additive manufacturing represents a cutting-edge technology that offers significant reductions in both manufacturing time and cost. However, any new technology or material must go through a qualification process before it can be used in the nuclear industry. This article reports on a pre-qualification process of 316L stainless steel manufactured using the laser powder bed fusion (LPBF) additive manufacturing process. The study compares LPBF 316L, aged, and non-aged materials from two different manufacturers. A small punch creep test campaign at 650 °C was performed at different loads. These tests are particularly advantageous because they require only a small amount of material, making them ideal when material availability is limited. Additionally, standard uniaxial creep tests were performed at the same temperature for comparative reference. A good correlation for the time to rupture–equivalent stress between the two test types was observed, with the equivalent stress calculated using the Small Punch Test EN 10371:2021 standard. A significant finding is that the small punch creep deflection rate curves for LPBF-manufactured 316L exhibit multiple minima, unlike the single minimum observed in forged 316L. This is believed to result from micro-cracking and has important implications for determining the equivalent stress creep properties, which are based on the single minimum value in the EN 10371:2021 standard. The multiple minima finding suggests that the approach used to determine equivalent stress and strain rate in small punch creep tests in the EN 10371:2021 standard must be re-evaluated to accommodate this complexity.

1 Introduction

Additive manufacturing (AM) represents a cutting-edge technology that potentially offers significant reductions in both manufacturing time and cost. Due to completely different principles of manufacturing, both the AM feed stock material and build process must be qualified, especially for application in the nuclear industry. The Procedure for the Acquisition of New Material Data (PTAN) of the RCC-MRx code for design and construction rules (AFCEN, 2017) and Section 3, Division 5 of ASME Boiler and Pressure Vessel Code (BPVC) (ASME, 2013) are two examples of procedures for the qualification of new materials. The qualification process includes defining a material property file, which, among others, includes physical and material properties from a wide range of tests and laboratories to establish material confidence limits in view of variation of properties. These can be due to material taken from a thick or thin plate, heat-to-heat variability, different cooling rates from AM melt, and later reheating by following passes of the heat source, etc.

The EU-funded research project NUclear COmponents Based on Additive Manufacturing (NUCOBAM) recently ended (2020–2024) (NUCOBAM, 2024). It examined the qualification of AM 316L stainless steel to demonstrate the compatibility of AM 316L in a light water reactor (LWR)-irradiated environment (Konstantinović et al., 2025). A more comprehensive program is currently being considered in ASME (Messner, 2023). In this work, laser powder bed fusion (LPBF) was used because no excessive ferrite phase builds up during rapid cooling.

The nuclear community (scientists, reactor designs, and regulators) is currently debating whether traditional material qualification tests, which assume that material properties derived from standard tests reduce uncertainty through extensive testing, are sufficient or if component-level tests are necessary (Messner, 2023; Torres and Gordon, 2021; Lalé and Viguier, 2024; ASM Handbook, 2017). In NUCOBAM, the process involved qualification of 316L feedstock powder first, followed by qualification of the AM build process and platforms, and qualification of the AM build material through a number of mechanical tests. Finally, a valve body and debris filter were manufactured for component testing. In this work, we focus on the recently standardized (EN, ASME, and RCC-MRx) small punch test (CEN EN, 2024; ASTM, 2024; RCC-MRx, 2022). This test uses small disks for test pieces (8 mm diameter, 0.5 ± 0.005 mm thickness), which is a crucial advantage when only a small amount of test material is available (Figure 1). Specifically, the small punch creep (SPC) tests were carried out to estimate creep properties, followed by validating these properties by performing a small number of standard uniaxial creep (UC) tests.

Figure 1

Figure 1. Uniaxial creep test piece (green) in relation to the small punch test piece given in two orientations (red and grey). The LPBF build direction is the Z-axis.

The article is structured as follows. We first describe the feedstock material, followed by a description of the methods used. Next, we discuss the specificities of the analysis of the SPC test results of AM 316L material and the transferability of these results to UC. Conclusions are given at the end.

2 Materials

A 316L stainless steel feedstock powder was used to manufacture bars for material tests using the laser powder bed fusion (LPBF) additive manufacturing process in accordance with the ASTM F3184 (ASTM F3184-16 Standard Specification for Additive Manufacturing Stainless Steel Alloy, 2024). The basis for the feedstock powder chemical composition requirements was taken from ASTM F3184, with additional requirements imposed for the weight % of N, O, Cu, Co, Ta, B, and Cu. The chemical composition of the feedstock powder is comparable to the 316L(N) forged plate but lower in the C (0.02 vs. 0.024) and Mn (0.8 vs 1.794) content and higher in Si (1 vs. 0.689). Table 1 provides the chemical composition, while the measured particle size distribution is given in Table 2.

Table 1

Table 1. Chemical composition of powder feedstock 316L stainless steel.

Table 2

Table 2. Measured size properties of powder feedstock 316L stainless steel.

The powder feedstock was then used to manufacture bars on a build platform (Figure 2). In this article, we use bars manufactured using a Renishaw AM250 single laser machine at the Advanced Manufacturing Centre (AMRC), University of Sheffield, and Commissariat à l'énergie atomique et aux énergies alternatives (CEA), using an SLM Solutions SLM280HL single laser, both using argon as an inert gas. The bars used in this work are in the as-built z-direction, protruding out of the build platform, as shown in Figure 2.

Figure 2

Figure 2. Build platforms: AMRC (left) and CEA (right).

A heat treatment consisting of soaking at 1,066 ± 14 °C with a holding time of 1 h 15 min ±15 min was applied to the AMRC build platform, while for the CEA, the bars were cut from the build platform prior to application of the same thermal treatment. The holding time counter initiated when a lower temperature limit, 1052 °C = 1066 °C–14 °C, was attained. An argon atmosphere was used, with the prescribed heating rate of 4 °C/min and a cooling rate of at least 10 °C/min. We refer to this heat treatment as HT2 to distinguish it from other heat treatments applied within NUCOBAM.

In addition to HT2, some bars were aged for 10,000 h at 450 °C. The following combinations of materials were used in the current article: a) AMRC, HT2; b) AMRC, HT2 aged; and c) CEA, HT2, aged. To simplify labeling, these will be referred to as AMRC NA (not aged), AMRC Aged, and CEA Aged. These materials were then used to manufacture all small punch and uniaxial creep test pieces.

Four bars from the internal AMRC qualification process were used for three UC tests (at initial stress of 250 MPa, 270 MPa, and 290 MPa) due to an insufficient amount of originally shipped material. The heat treatment of these bars was performed at the JRC, following the same HT2 procedure.

3 Methods

3.1 Small punch creep tests

A two-stage approach was taken for estimating creep properties. In the first stage, a small punch creep (SPC) test campaign was performed at 650 °C. The SPC test involves pressing a 2.5-mm-diameter punch (a high precision silicon nitride Si₃N₄ bearing ball) into a test piece at a constant force (Figure 3). The test piece is a disk with a diameter of 8 mm and a thickness of 0.5 ± 0.005 mm. The test piece is clamped between the upper and lower dies. The lower die has a 4-mm-diameter receiving hole with a 0.2 mm/45° chamfer. The deflection of the test piece is measured from below the test piece with a ceramic rod touching the test piece and extending to a linear variable differential transformer (LVDT) gauge. The LVDT has a negligible spring counter force. The result of an SPC test is a deflection versus time curve, resulting in a given time to rupture, t_r.

Figure 3

Figure 3. Illustration of the small punch test.

Small punch creep (SPC) tests were done at 650 °C with the test piece in an argon environment to prevent test piece corrosion. Force levels from 240 N to 500 N were used by the application of dead weights. The test matrix is given in Table 3. This test matrix was then used to estimate the equivalent load conditions of uniaxial creep tests as described below.

Table 3

Table 3. SPC test matrix at 650 °C.

From the SPC deflection versus time curve, one can estimate the equivalent initial stress, ${\sigma }_{eq}$, to apply to a UC test (at the same temperature) that results in the same rupture time. The informative Annex G in EN 10371:2021 standard (CEN EN, 2024) provides Equation 1 for estimating ${\sigma }_{eq}$. We distinguish here between ${\sigma }_{eq}$, calculated using Equation 1 from the SPC test, and ${\sigma }_{init}$, which is the initial stress applied to the uniaxial creep test.

$\frac{F}{{\sigma }_{eq}}=1.916·u_{\mathit{min}}^{0.6579}\left[\frac{\mathrm{N}}{\text{MPa}}\right](1)$

In Equation 1, F stands for the force applied in the SPC test, and $u_{\mathit{min}}$ stands for the deflection in the SPC test where the minimum deflection rate, $\dot{u}$, is obtained.

3.1.1 Estimation of equivalent stress ${\mathrm{\sigma }}_{eq}$

Accurate deflection measurements and estimation of deflection rate are essential for calculating equivalent stress, Equation 1, and creep assessment in general. Temperature control precision, susceptibility of the laboratory environment to external vibrations, and systematic noise are only a few of the influencing parameters. A typical recorded deflection during an SPC test in our laboratory is given in Figure 4. A sampling rate of 1 Hz is used to capture the initial load application. Because our data acquisition software can only sample at a constant rate, a 1 Hz sampling rate is maintained for the duration of the test.

Figure 4

Figure 4. A typical deflection in an SPC test. Insert: zoom in on the measured deflection (noise).

The deflection signal contains approximately ±1 to ±2 µm noise (Figure 4 insert). It is well known that differentiating a signal increases the noise. It is clear that such a signal cannot be used for evaluating equivalent stress, Equation 1, where the deflection at a minimum deflection rate must be identified.

We tackle the issue by first down sampling the deflection signal to 2¹⁴ = 16,384 points using a linear interpolation. This significantly reduces the number of points for processing the signal, decreasing computational time for noise filtering and calculation of deflection rate. A large number of points can result in a prohibitively large amount of required memory. A larger number of points can be used for longer experiments. In the second step, we deal with the noise. One way is to split the signal into segments and calculate the linear regression line for each segment. We investigated the impact of 50, 100, 200, and 300 segments. Using 200 segments was determined to be the optimal number in terms of minimizing noise while still retaining sufficient local detail (important for, e.g., determining the crack initiation). The deflection rate is then the inclination of the regression line. This decreases noise in the calculated deflection rate, but significant local peaks remain; see the black solid line in Figure 5. If a Whittaker noise filter (Eilers, 2003) is applied prior to calculating the regression lines, the noise in the calculated deflection rate improves further. However, there are still a number of local variations; see the red solid line in Figure 5. We also tried several methods for calculating a derivative, available in the Python “derivative” library. The best performing method was using a “spline” derivative; see the solid green line in Figure 5. The combination of a Whittaker noise filter and the spline derivative method consistently resulted in very good noise removal, preservation of significant local variations in the deflection curve (see the red circle), and few artificial peaks in the calculated deflection rate. This combination is therefore used in all further results in this article.

Figure 5

Figure 5. Deflection rate calculation methods.

3.2 Uniaxial creep tests

In the second stage, reference uniaxial creep tests were performed at 650 °C. The equivalent stress ${\mathrm{\sigma }}_{eq}$ was estimated from SPC tests, using Equation 1, and then initial (uniaxial) stress levels between 150 MPa and 290 MPa were selected (Table 4).

Table 4

Table 4. UC test matrix at 650 °C.

4 Results and discussion

Figure 6 shows the SPC initial deflections, u₀, defined as a deflection 5 s after the application of the full force load. These deflections are the result of an elastic–plastic response of a test piece. Because the calculation of the equivalent stress ${\mathrm{\sigma }}_{eq}$, Equation 1, depends strongly on the deflection, it is important to accurately capture these deflections. The initial deflections follow a linear trend relative to the load, with some scatter even within the same material. AMRC NA tests were performed first and have the largest scatter. Lessons learned in applying the test procedure reduced the scatter later on during AMRC Aged and CEA Aged tests. All the test pieces conformed to the very strict EN 10371:2021 thickness tolerances (0.5 ± 0.005 mm), and both the test piece temperature and the force load were strictly controlled. Some scatter in results within the same material is probably due to differences in the microstructure of test pieces. Because LPBF is used, the test pieces are basically welds, and local variations in material properties can be expected. Furthermore, surface roughness of the test pieces and the silicon nitride Si3N4 bearing ball, resulting in different friction effects between the tests, could be a contributing factor. The minor influencing factor could be the small variations in the silicon nitride Si₃N₄ bearing ball diameter (grade 5, ±0.0013 mm diameter tolerance) and load application rate, which we are not able to strictly control on our test machines.

Figure 6

Figure 6. Measured initial elastic–plastic deflection u₀.

Figures 7–14 show measured deflections during SPC tests at load rates from 500 N to 240 N (Jorge and Myriam, 2025). The symbols indicate the time at which the minimum deflection rate was estimated, using the procedure described above. In general, the aged material exhibits shorter creep life. In addition, the minimum deflection rate occurs earlier for the aged material; see Table 5. The exception is the AMRC NA FK-13 test at 260 N, where the minimum deflection rate occurs earlier than in both CEA Aged tests (GB-08, GB-09). In general, the CEA Aged creep lives are somewhat shorter than AMRC Aged lives at loads from 500 N to 300 N. At 260 N and 240 N loads, there is no significant difference between the two.

Figure 7

Figure 7. Measured deflection and calculated deflection rates at the 500 N load level.

Figure 8

Figure 8. Measured deflection and calculated deflection rates at the 450 N load level.

Figure 9

Figure 9. Measured deflection and calculated deflection rates at the 400 N load level.

Figure 10

Figure 10. Measured deflection and calculated deflection rates at the 350 N load level.

Figure 11

Figure 11. Measured deflection and calculated deflection rates at the 320 N load level.

Figure 12

Figure 12. Measured deflection and calculated deflection rates at the 300 N load level.

Figure 13

Figure 13. Measured deflection and calculated deflection rates at the 260 N load level.

Figure 14

Figure 14. Measured deflection and calculated deflection rates at the 240 N load level.

Table 5

Table 5. SPC test results.

In Figure 9, we can see that all the deflection rate curves at 400 N force load exhibit two minima, with the first one having a significantly lower deflection rate than the second one. The two minima are even more apparent at decreased force load, 350 N, but are now almost evenly matched in terms of deflection rate. Calculation of the equivalent initial stress ${\mathrm{\sigma }}_{eq}$, Equation 1, according to the EN 10371:2021, depends on the identification of the minimum deflection rate. However, the standard assumes that only one minimum deflection rate is observed. At 350 N, it is unclear if we should select the first or the second minimum, especially because they are evenly matched in terms of deflection rate, and the estimation of the deflection rate itself is subject to the approach for dealing with the noise in the measured deflection signal and the derivative calculation method.

Another question that emerges is what causes the first minimum. In this work, we conjecture that a significant crack develops and manifests in the first minimum. It is well known that the region of maximal stress/strain in small punch tests continuously moves radially outwards during the test. We therefore conjecture that if the material is ductile enough, the crack that results in the first minimum in the deflection rate curve arrests as the zone of maximal stress/strain redistributes radially outwards. At a certain point, a second crack initializes and starts to propagate, resulting in the second minimum, shortly before the rupture of the test piece. At the 240 N load, one can see a significant local increase in the deflection at approximately 0.9 mm for the FK-10 AMRC NA and FV-08 AMRC tests (the orange and green curves in Figure 14). We assume these sudden deflection increases are due to a final/secondary crack initialization. To evaluate these conjectures, we performed an interrupted SPC test with AMRC NA material. Because no more test pieces conformant with the strict 0.5 ± 0.005 mm requirement of the EN 10371:2021 were available, we used a test piece with a thickness of 0.4661 mm, labeled FI-07. The EN 10371:2021 standard provides a formula for estimating the ultimate tensile stress, $R_m$, from the small punch tensile test (Equation 2). In the formula, the ultimate tensile stress is linearly related to the initial thickness of the test piece, $h_0$. ${\beta }_{Rm}$ stands for the correlation factor, $F_m$ is the maximum force, and $u_m$ is the deflection at maximum force. Assuming that $u_m$ is approximately the same, we can calculate the force for the thinner test piece that would be equivalent to the 0.5-mm test piece (Equation 3). Recognizing that this is a very crude assumption, we calculate the force to be applied as 326 N = 350 N × 0.4661 mm/0.5 mm.

$R_m={\beta }_{Rm}·F_m/\left(h_0·u_m\right)(2)$

$F_{mthin}=F_m·\frac{h_{0thin}}{h_0}(3)$

Based on the SPC tests at 350 N, we decided to interrupt the SPC test for the first time 12 h after application of the force load of 326 N. We defined the interruption as stopping the furnace heating by setting the target furnace temperature to 22 °C, triggering the cool down. During the cool-down, the force load remained at 326 N. Following the cool-down, we removed the dead weights, removed the test piece from the test machine, and performed a scanning electron microscope (SEM) analysis of the test piece. Two crack fronts are visible, as shown in Figures 15, 16. The first crack front with a diameter of 1.28 mm is fully developed after 12 h. The second front, with a diameter of 1.46 mm, has only started developing.

Figure 15

Figure 15. SPC interrupted test, FI-07 test piece, 326 N load level. Crack fronts 12 h after the start of the SPC test.

Figure 16

Figure 16. SPC interrupted test, FI-07 test piece, 326 N load level. Crack fronts 12 h (top) and 28 h 48 min (bottom) after the start of the SPC test.

Next, we put the FI-07 test piece back into the test machine, added the same dead weights to reapply a 326 N force load, and heated the test piece back to 650 °C. After reaching the target temperature, we waited for 16 h and 48 min (cumulative time 12 h + 16 h + 48 min) and then initiated the second interruption, using the same procedure as for the first interruption. A second SEM analysis was performed (see the lower part of Figure 16). The analysis showed the arrest of the first crack front and full development of the second, much larger crack front that only started forming after the first interruption. We did not continue the SPC tests beyond this point, 16 h and 48 min (cumulative time 12 h + 16 h + 48 min). The deflection versus time curve and the calculated deflection rates indeed show two minima (Figure 17), although these minima are less pronounced than the ones at 350 N (Figure 10). Nevertheless, this confirms our conjecture that cracks formed during the SPC test. More interrupted tests are required to confirm that crack initiation results in a local minimum in the deflection rate.

Figure 17

Figure 17. Interrupted AMRC NA test. Deflection was measured and deflection rates were calculated at the 326 N load level.

Figure 18 shows the time to rupture versus the applied force for the SPC test data for the three materials and the regression lines. This confirms again that the aged materials have consistently reduced creep life compared to the non-aged material.

Figure 18

Figure 18. SPC: measured time to rupture versus force.

Nuclear design codes, and in particular evaluation of material properties, are primarily based on standard uniaxial tests. Uniaxial tests were conducted to assess the SPC test data and to assess the transferability. To this end, we calculated the equivalent stress ${\mathrm{\sigma }}_{eq}$, Equation 1, for all the SPC tests, and compared ${\mathrm{\sigma }}_{eq}$ versus rupture time to the applied initial stress of UC tests ${\mathrm{\sigma }}_{init}$ versus rupture time, Figure 19. One UC test was performed for each material at 150 MPa, 170 MPa, and 190 MPa initial stress ${\mathrm{\sigma }}_{init}$. For the AMRC NA, we also performed UC tests at 210 MPa, 250 MPa, 270 MPa, and 290 MPa initial stress ${\mathrm{\sigma }}_{init}$. Table 6 gives an overview of the UC test results.

Figure 19

Figure 19. SPC and UC: measured time to rupture versus ${\sigma }_{eq}$ (Equation 1), ${\sigma }_{init}$.

Table 6

Table 6. UC test results.

We can observe that UC tests for the three materials overlap perfectly and are distributed along the regression line with an R² = 0.9938, Figure 19. In contrast to SPC tests, there is practically no difference between the three materials in UC tests. Furthermore, the scatter in the UC results between three different materials is much smaller than in the SPC results. We can also see that the transferability of SPC results to UC, by evaluating ${\mathrm{\sigma }}_{eq}$ from the SPC tests, is relatively good. Equation 1 from the EN 10371:2021 standard works quite well. One can therefore estimate the equivalent UC initial stress load condition from the SCP tests. However, the evaluated ${\mathrm{\sigma }}_{eq}$ slightly overestimates the stress below t_r = 50 h and underpredicts it above that value due to different slopes of the SPC and UC trendlines.

Several explanations are possible for the SPC trendline having a different slope than the UC trendline. First, Equation 1 is based on one minimum in the SPC deflection rate, whereas here, we can have two minima, as shown in Figure 10. In this work, we always select the one with the lowest deflection rate value to use in Equation 1. However, as shown in Figures 15–17, the first minimum is probably the one where the initial crack has already formed. If the second minimum is used, the measured deflection is not solely due to creep but is significantly affected by the crack propagation. Second, the orientation of the SPC test pieces in relation to the uniaxial test piece is the one shown in red in Figure 1. This was driven by the limited amount of material, as more disc slices for SPC test pieces can be extracted from a bar in the red orientation than in the gray SPC orientation. In the SPC test, most of the stress acts in the radial direction of the SPC test piece. With the red SPC test piece orientation, the red SPC test piece has stress perpendicular to the stress in the uniaxial test piece. We are therefore testing a different material direction in the SPC tests. In fact, we should have used the gray SPC test piece orientation, but as mentioned, this was not possible due to the limited amount of available material. The SPC test piece orientation in relation to the test build can have a significant impact on the time to rupture in LPBF. For Inconel 718, Zhang et al. (2025) report significantly shorter rupture times for the SPC test piece in the build direction (red color in Figure 1) compared to the one perpendicular to the build direction (gray color in Figure 1). In this case, only aged heat treatment was applied. However, when they applied the solution and aged heat treatments, the differences in rupture times decreased significantly. This behavior is at least in part driven by a different microstructure (Zhang et al., 2025) of the two directions. Third, and probably the most important, Equation 1 has been developed based on 97 UC tests and 159 SPC tests with both new and service-exposed materials, mainly low-alloy steels and 9Cr steels, such as 14MoV63, X20CrMoV121, P91, P92, Eurofer97, and stainless steel 316L (CEN EN, 2024). However, this data set had a large share of P92 steel. Equation 1 coefficients 1.916 and 0.6579 are probably not the best for AM 316L material.

4.1 Microstructural analysis

The fracture surfaces of AMRC NA, AMRC Aged, and CEA Aged UC test pieces, at a load level of ${\mathrm{\sigma }}_{init}$ = 150 MPa, were analyzed by scanning electron microscope (SEM). Figures 20, 21 show similar patterns obtained for both the AMRC NA and CEA Aged test pieces, respectively. The surfaces are not flat and exhibit a mixed ductile-brittle fracture. For the same test piece, one of the fracture surfaces exhibits cavities while the other shows slight spherical extrusions, roughly 50 µm in diameter. The parallel patterns observed in the structure are likely due to the laser scanning paths followed, related to scanning speed and hatch spacing. The fracture surface morphologies suggest that the overlapping regions of two laser tracks generate more ductile thin areas where some dimples can be distinguished. On the other hand, the cavities and elevation exhibit a more brittle aspect, with some cleavage features.

Figure 20

Figure 20. SEM picture of both UC fracture surfaces in the XY orientation. AMRC NA, test piece FA-01, ${\sigma }_{init}$ = 150 MPa.

Figure 21

Figure 21. SEM picture of both UC fracture surfaces in XY orientation. CEA Aged, test piece GC-01, ${\sigma }_{init}$ = 150 MPa.

5 Conclusion

In this article, we provide a significant set of small punch creep (SPC) and uniaxial creep (UC) data of additively manufactured (AM) 316L stainless steel at 650 °C. We show that SPC tests of this material result in double minima for deflection rates at relatively short rupture times (tens of hours). For a specific case, we show that these two minima are related to crack initiation and propagation. The measured SPC deflection is then not only due to the creep of the test piece but also due to its cracking. This could potentially be an issue for SPC testing of non-ductile materials. Further work is necessary to confirm additionally that the multiple crack initiations in SPC tests result in multiple deflection rate minima, to determine which minimum should be used for estimating ${\sigma }_{eq}$, and/or whether a completely different method should be used. We demonstrate the transferability of SPC results to UC results by calculating that the ${\sigma }_{eq}$ works quite well for the AM 316L in this work. However, the ${\sigma }_{eq}$ underestimates the stress for the shorter rupture times and overestimates the stress for longer rupture times. This is probably due to the EN 10371:2021 standard assuming a single minimum in the SPC deflection rate, while we show here that two minima can occur. It is clear, however, that the EN 10371:2021 standard must be re-evaluated to accommodate this complexity.

Data availability statement

The datasets presented in this study can be found in online repositories. The names of the repository/repositories and accession number(s) can be found below: NUCOBAM SPC, 650C catalogue, https://doi.org/10.5290/76, NUCOBAM UC, 650C catalogue, https://doi.org/10.5290/77.

Author contributions

IS: Conceptualization, Data curation, Formal Analysis, Investigation, Methodology, Project administration, Resources, Software, Supervision, Validation, Visualization, Writing – original draft, Writing – review and editing. KN: Formal Analysis, Methodology, Writing – original draft, Writing – review and editing. SH: Conceptualization, Methodology, Writing – review and editing. MK: Investigation, Writing – review and editing. AG-J: Investigation, Writing – review and editing.

Funding

The author(s) declare that financial support was received for the research and/or publication of this article. This project has received funding from the European research and training programme 2014–2018 under grant agreement No. 945313 (NUCOBAM project).

Conflict of interest

The authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.

Generative AI statement

The author(s) declare that no Generative AI was used in the creation of this manuscript.

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