Figure 4.5 illustrates the calculation of the NZPV using the thermal imaging method. The position of the magnetic defect is known on the recording and hence the centre line from where the normal zone propagates x0 as well. On a subsequent image at time i the propagation of the 93 K zone can be observed upwards until point xi up. From this, the NZPV was calculated as

$$\text{NZPV}\_{\text{up}}^i = \frac{\Delta x}{\Delta t} = \frac{x\_{\text{up}}^i - x\_0}{t\_i - t\_0},\tag{4.2}$$

where Δx and Δt represent the spatial and temporal change in the propagation, respectively. If a visible difference was observed in the travelled distance by the normal zone in the opposite direction (downwards in this case), the NZPV was calculated at a later time i + 1 as

$$\text{NZPV}\_{\text{down}}^{i+1} = \frac{\Delta x}{\Delta t} = \frac{x\_{\text{down}}^{i+1} - x\_{\text{down}}^{i}}{t\_{i+1} - t\_{i}}.\tag{4.3}$$

1 Common values used for the Tc of HTS coated conductors is in the range of 90 K to 93 K [6, 48, 73, 108, 109].

Figure 4.5: Normal zone propagation in HTS tapes showing the steady-state operation at 77 K as well as two time instances after the appearance of a normal zone. The red coloured area indicates the propagation of the 93 K front, used as a criterion for calculating the NZPV.

An actual example of a post-processed recording is shown in Figure 4.6, with the visual aid of the overlaid gridlines in 2 mm intervals. In this example the centre line, where the normal zone started developing, was at x0 = 42 mm, as measured from the upper part of the tape. The normal zone development itself was first observed at approximately t0 = 7 ms. Starting from this time and position, Figure 4.6a indicates that the normal zone propagated 4 mm up and 4 mm down from the nucleation point until the 12.57 ms mark. Then, an additional 2.57 ms later in Figure 4.6b (absolute time of 15.14 ms), the normal zone travelled 6 mm in both directions. By knowing both the elapsed time and distance travelled (with respect to t0 and x0, respectively), the NZPV was obtained using eq. 4.2 for the first time step

$$\text{NZPV} = \frac{0.004\,\text{m}}{0.01257\,\text{s} - 0.007\,\text{s}} = 0.72\,\text{m}\,\text{s}^{-1} \tag{4.4}$$

and eq. 4.3 for the second time step

$$\text{NZPV} = \frac{0.002\,\text{m}}{0.015\,14\,\text{s} - 0.012\,57\,\text{s}} = 0.78\,\text{m}\,\text{s}^{-1}.\tag{4.5}$$

Figure 4.6: Example of a recording post-processed for calculating the normal zone propagation velocity. (a) shows the normal zone at time 12.57 ms and (b) at 15.14 ms. Both show the raw frame (left) and a thermal image (right), with the scale boundaries between 77 K and 93 K. The red grid lines are 2 mm apart and the arrows indicate the nucleation point (caused by the magnetic defect) and the distance travelled by the quench in the up- and downwards direction.

Images processed in this manner are difficult to use for other purposes and hence this post-processing was solely used as an intermediate step for calculating the thermal NZPV. For general visualization purposes and observing the behaviour of the quench the grid lines were not plotted.

It must be noted that – especially in the case of some 12 mm-wide tapes – determining an accurate NZPV using the above described method became tedious. It happened on occasion that the 93 K wave front did not propagate uniformly; one side of the tape quenched more rapidly than the other side, for example. In such situations it became difficult to determine if the tape really quenched or not or to what extent. To avoid such problems, in the measurements the criterion was further specified arbitrarily as "the furthest point of the 93 K zone". Apart from this, the boiling also severely distorted the measurements and the temperature rises were on occasion so quick that defining a 93 K wave front did not make a practical sense. This happened frequently in stabilizer-free tapes during a quench, where the tape was locally often either at 77 K or at some higher temperature, above the measurement limit of 260 K of the thermal imaging method.

The method is further limited in measuring the NZPV in the case of copper-stabilized HTS tapes under certain current conditions. Exceeding the critical current of the tape caused on more than one occasion the current to transfer into the copper shunt practically instantaneously. While this is the expected behaviour and exact function of the shunt layer in the first place, in such situations the HTS tapes heated almost uniformly over the length. On other occasions (mostly stabilizer-free) tapes rapidly developed film boiling over the entire surface, which then distorted the recording and as a consequence, no "thermal wave front" was visible that would allow calculating the NZPV. This was also due to current bypass into the thin silver layer, however due to the significantly lower heat capacity, the layer quickly heated up and entered a film boiling regime (refer to Figure A.3 in the Appendix). An alternative here can be the use of the boiling development over the tape's surface, however, it is unknown how this then compared to the measured NZPV as described previously.

For these reasons, while with electrical NZPV measurements it is possible to measure the quench propagation velocity at multiple current levels (even far above the tape's Ic), this is most often not the case with the described thermal imaging method.

## 4.2.2 Results and comparison

The results of the NZPV measurements are summarized in Table 4.1 together with the relevant properties of the tapes, where the normal zone propagation velocity could be reliably calculated from the thermal imaging as introduced. The discussion about the results is subdivided into three main groups based on the architecture of the tapes: stabilizer-free HTS tapes, HTS tapes with metallic stabilization and HTS tapes with some form of special architecture.

### HTS tapes without stabilization

Results of sample 1 from Theva showed a good correlation between the thermal imaging and the electrical measurement. With the tape's critical current in the range of 600 A to 700 A, 650 A and 625 A pulses for a duration of 12 ms have both recorded a normal zone propagation velocity of 5 m s−1.

Sample 2 was practically identical to sample 1, however, this was damaged during the pulsed measurements at KIT. Here a 5-fold lower NZPV was measured at merely 63 % of the critical current.

Sample 6 was a tape from SuperPower with only an Ag-layer. With a current pulse amplitude of 158 A, an NZPV of 1.82 m s−1 was measured electrically. With the thermal imaging method a normal zone propagation was only observed much higher at 180 A (over 60 % above the critical current) with a mere 0.83 m s−1. This is about half that of the electrical measurement result and the exact reason is unknown.

A similar mismatch was observed in the electrical and thermal measurement results of sample 7 from SuperPower. Pulsing the tape with 100 % of its critical current resulted in an NZPV of 1.96 m s−1 in the electrical measurement and only 0.17 m s−1 in the thermal imaging. This constitutes an order of magnitude difference in the two measurement methods.

Sample 4 from SuperPower was interesting, since while with the electrical measurement method an NZPV was determined for a current pulse of 85 A as 0.3 m s−1, at this current level the tape did not show any signs of a quench when measured at KIT. A quench was observed only at 115 A first with a normal zone propagation velocity of 0.71 m s−1.

## HTS tapes with stabilization

Sample 3 was a Theva sample with a nickel-copper alloy as a stabilization. The NZPV was measured as 0.5 m s−1 at both institutes when pulsed with 600 A (Ic), showing a close match between the two methods.

Sample 9 was a sample from SuperOx with a standard 20 μm copper stabilization. Applying a current pulse 20 % above the tape's Ic resulted in an NZPV of 0.5 m s−1, however, data from Montréal was unfortunately not available for this tape.

#### HTS tapes with special architecture

Two tapes with special, current flow diverter architectures [110] were also measured (this type of tape architecture is introduced in greater detail in Section 4.3.3). Sample 5 was prepared from a sample practically identical to sample 4 by the group at Montréal. While the electrical NZPV measurement showed a quench propagation speed of 3.3 m s−1, an over 7-fold improvement, with respect to sample 4, thermal imaging could not reliably determine an NZPV. In the measurements, the tape quickly developed a boiling region over the entire surface and as a consequence no travelling wave front was observed. It is believed, however, that this is the sign of a considerably improved NZPV as previously discussed in Section 4.2.1.

Another current flow diverter sample was prepared by Oxolutia with a novel process using inkjet printing [111, 112]. A sample was measured successfully – for the first time – at KIT for NZPV with the fluorescent thermal imaging method. A speed of around 7.5 m s−1 was obtained at approximately 93 % of the critical current. This sample is discussed in more detail in Section 4.3.

## 4.3 Qualitative quench investigation

The promise of high-speed fluorescent thermal imaging is the additional, qualitative information provided about the quench development aside the capability of calculating the NZPV. This aspect of quench is not often researched and therefore constitutes the focus point of this section.

Here the samples 3, 4, 5, 8 and 9 from Table 4.1 are discussed in greater detail with respect to the quench formation and development. Electrical measurements are shown for different current pulses together with both thermal and raw fluorescent images. These latter are often valuable in recordings where liquid nitrogen boiling considerably distorted the image. In such cases, the quench formation is better visible on the raw images than on the post-processed thermal images.

## 4.3.1 HTS tape without stabilisation

Here the electrical and thermal imaging results of a 4 mm-wide, 120 mm-long HTS sample from SuperPower are presented on the example of sample 4 in Figure 4.7. The stabilizer-free sample had solely a 1.2 μm thick Ag coating, making it ideal to show a characteristic quench propagation of such a tape. The Ic of the sample was 85 A, however, at current pulses up to 110 A no quench was visible on the thermal imaging.

The first six current pulses in Figure 4.7 show clearly visible voltage signals, but again, no quench on the thermal imaging. The pulse duration was kept constant at 10 ms and the current amplitude was increased from 85 A to 110 A. The absorbed energies

Figure 4.7: Sample 4 pulsed with 85 A to 110 A, 10 ms long pulses. Clear voltage signals are visible, however, the thermal imaging showed no signs of a quench.

were 1.53 J, 1.62 J, 1.71 J, 1.8 J, 1.87 J and 1.98 J, respectively, calculated from eq. 4.1. Electric fields of up to 9Vm−1 were recorded that are in the range where the electrical signals are used for NZPV calculation.

The current amplitude was increased further and a quench became first visible at 115 A (135 % Ic). As indicated in Figure 4.8a, on the thermal imaging the quench is first visible at 12.65 ms where the measured electric field was 6.3 V m−1. The extracted images in Figure 4.8b-4.8d show uniform quench propagation along the width of the tape and eventually heating to beyond 260 K. Note the gradual voltage development across the sample over the time of the pulse. Calculating the total deposited energy based on eq. 4.1 results in 4.36 J, whereas the absorbed energy until a visible quench was merely 0.57 J.

Figure 4.8: Sample 4 pulsed with a 115 A, 12 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a).

## 4.3.2 HTS tapes with stabilisation

## Theva tape with 100 μm Cu-Ni

Sample number 3 was prepared by Theva, using a special, 100 μm copper-nickel electrical shunt (refer to Table 4.1) and the measurements on this sample show well the general behaviour of metal-stabilized tapes. In the following, three separate measurements are discussed where different current pulses were applied and, accordingly, the tape quenched in different ways. These were the following

The shortcomings of the thermal imaging method are clearly visible – as it often happened with metal-stabilised samples – where the current bypass rapidly heated up the tape surface and formed a film boiling layer. This layer then partially or completely prevented obtaining temperatures over the surface, however, the raw images still showed the general quench behaviour.

Figure 4.9 shows the sample's behaviour when pulsed with the 700 A, 12 ms long pulse. Unfortunately, the quench is only visible on the raw images. Here a uniformly propagating quench region is visible, starting at the magnetic defect's location. It is also clear that the quench has only propagated about ∼0.5 cm in total as shown in Figure 4.9d. The energy deposited into the tape was also calculated according to eq. 4.1, which yielded 2.09 J.

The sample was subsequently pulsed with 900 A for the same 12 ms duration as indicated in Figure 4.10a. At this current level, the voltage drop across the sample became too high and the current source went into voltage limiting, causing a drop in the output current. Nevertheless, after the initial quench zone formation at the magnet's position, several additional quench zones became visible, as shown in Figure 4.10d. The appearance of multiple quench zones – here and in other measurement as well – is hypothesised by the presence of material defects, causing local drops in the tape's Jc and thereby facilitating further quench nucleation after the initial transition began. The energy deposited in the tape during this pulse was 19.64 J, over the same 12 ms pulse duration.

In a third measurement a lower, 600 A pulse was applied but for longer, 100 ms. As shown in Figure 4.11a, the current source was yet again driven into voltage limit and the current across the sample dropped. However, this time the sample only quenched at the magnetic defect and the quench propagated from there. No multiple quench zones

Figure 4.9: Sample 3 pulsed with a 700 A, 12 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a).

were observed in this experiment and the thermal imaging showed heating to over 260 K, due to an overall deposited energy of 51.42 J. This energy is over twice as much as in the previous experiment, however it was absorbed over the duration of 100 ms.

The three measurements would seem to indicate that a current pulse close to the Ic of the conductor would create a quench at the weakest point in the tape, at the magnetic defect in this case. If the pulse duration is short, the normal zone simply does not have enough energy (or time) to propagate. Increasing the pulse duration provides this additional energy and hence causes a propagating quench zone. On the other hand, if a current pulse considerably above the tape's critical current were applied, quench zones appeared at all (hypothesised) defects in the conductor and spread the energy out over the entire surface.

Figure 4.10: Sample 3 pulsed with a 900 A, 12 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a). The red arrows indicate the quench nucleation positions.

Figure 4.11: Sample 3 pulsed with a 600 A, 100 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a).

## SuperOx tape with 20 μm Cu

Similar behaviour was observed on sample 9 from SuperOx with a 20 μm pure copper stabilization. The tape was pulsed with

Figure 4.12 shows the 500 A pulse applied to the tape for 50 ms. While initially a small quench zone appeared, Figure 4.12b, quickly multiple quench zones developed as shown in Figures 4.12c and 4.12d. This is again hypothesised to happen since the tape was pulsed with a high, 1.5Ic, pulse. Due to the longer, 50 ms pulse duration all of the initiated quench zones began spreading and formed a film boiling over the surface. This again indicates that current pulses above the conductor's critical current cause quench nucleation at several locations, possibly material defects. Calculating the total energy deposited in the tape during the pulse shows that 8.45 J were absorbed.

In the second measurement the tape was pulsed with an even higher amplitude, but for a shorter duration, 700 A for 12 ms. Here the deposited energy during the pulse was 7.02 J, about 15 % less than in the previous measurement. However, as shown in Figure 4.13, the tape behaved completely differently. Again, the shortcoming of the thermal imaging method is obvious: it is impossible to obtain temperatures during the quench as the heavy surface boiling completely distorts the image. Regardless, the raw fluorescent images can be relied on. Only a single quench zone appeared, and quickly thereafter the entire surface of the tape developed a film boiling. It is furthermore interesting – although difficult to see on static images – that even though the quench initially started at the magnetic defect, it quickly distributed along the edges over the length of the tape and thereafter formed the film boiling. Applying currents far above the critical current caused the current to transfer rapidly in the metallic shunt, where due to its finite resistance this heated up, and caused surface boiling over the entire surface of the tape, distributing the energy.

Figure 4.12: Sample 9 pulsed with a 500 A, 50 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. The arrows show the point where the quench is first visible on the thermal imaging recording and where multiple quench zones started forming. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a). The red arrows indicate the quench nucleation positions.

Figure 4.13: Sample 9 pulsed with a 700 A, 12 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. The arrows show the point where the quench is first visible on the thermal imaging recording and where multiple quench zones started forming. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a). The red arrows indicate the quench development.

## 4.3.3 Tapes with current flow diverter architecture

A special application of fluorescent thermal imaging can be the qualitative inspection of quench in HTS tapes with special or experimental architectures, aimed at e.g. accelerating the NZPV. A so-called "current flow diverter" (CFD) architecture was proposed by Lacroix et al. [110] with this exact purpose. The principle is that by creating an area of high inter-facial resistance between the superconducting layer and the upper silver stabilisation along the centre of the tape, see Figure 4.14b, in case of a quench the current is forced to flow along the edges, a lower resistivity path. The increased current density in this smaller area causes local heating and propels the quench quickly along the edges.

An alternative implementation of the CFD architecture is called b-CFD [47], where no high-resistivity buffer layer is necessary between the superconducting layer and the Ag coating (Figure 4.14c). Instead, the thickness of the silver stabilization on the surface of the tape (superconductor side) is reduced, and simultaneously increased at the back (Hastelloy® side). Then by adding a thin conducting bypass around the longitudinal edges of the tape (e.g. silver) the current during quench can be forced to transfer via the edges to the bottom of the tape, a path with lower resistance. This, in effect, achieves the same goal of accelerating the quench propagation via the low heat capacity electrical conductor along the edges of the tape.

Figure 4.14: Cross-section of various HTS tape architectures. Subfigure (a) shows a common HTS tape, with uniform inter-facial resistance, (b) shows the CFD architecture with an artificially increased inter-facial resistance and (c) shows the b-CFD concept where the thickness of the Ag stabilization is changed on the front- and back-side of the tape [47]. The arrows indicate current sharing paths.

tecture.

#### Inkjet printed CFD

Oxolutia is a company based in Barcelona, Spain, developing a novel method to introduce a CFD architecture in HTS tapes. Their patented approach [113] is to use inkjet printing to apply a highly resistive interface layer between the superconductor layer and the silver finish [111, 112]. A sample was prepared by Oxolutia, using a Theva tape as a basis, with a 100 μm thick substrate. On top, 3.5 μm MgO and a 3.1 nm thick GDBCO layer were deposited. This was then followed by inkjet printing a ∼110 nm thick Y2O3 layer over 90 % of the width (centred), leaving the sides uncovered. The tape was then stabilized by sputtering 1 μm Ag on the surface. As a result, the tape edges have a low contact resistance between the superconducting and the Ag layer. Simultaneously, the Y2O3-covered part of the tape has a high interface resistance and achieves the CFD architecture similarly as shown in Figure 4.14b. Several shorter, 6 cm, samples were prepared with the above manufacturing method and were tested at KIT to determine both the NZPV as well as to observe whether the tape behaves differently during quench. Unfortunately, the high interface resistance – together with the short length – caused problems when injecting current in the samples. Often the contact resistance was either too high to force current in the sample with the available current supply or on other occasions the superconductor layer burned out under the current contacts. No reliable measurements could be done on such short samples. The first successful measurement was carried out on a longer, 10 cm sample. The Ic distribution of this sample was measured at the Institut de Ciència de Materials de Barcelona (ICMAB) using Scanning Hall Probe Microscopy [114]. The sample was found to have an Ic of over 700 A in a magnetic field of 47.9 mT.

Since this was a prototype HTS sample without any metallic stabilization, 12 ms current pulses were applied where the current amplitude was gradually increased. The sample eventually did quench at a current amplitude of 650 A, see Figure 4.15a, and a peak NZPV of 7.5 m s−1 was measured. The reason it is referred to as "peak NZPV" is since the current got limited in the sample at around 11 ms, after which the quench propagation slowed down. The electric field appeared suddenly across the sample and saturated the amplifier's 1 V range. However, because the current across the sample dropped, the current source must have gone into voltage limiting due to the high resistance across the sample, which happens at ∼10 V. Over the distance of the voltage taps in this measurement, the electric field across the sample was, at least, 179 V m−1. The appearance of the voltage signal matched exactly the visual appearance of a quench in the sample.

The sample also revealed an interestingly-shaped quench zone. Looking at Figure 4.15c a hot zone is visible around the centre of the tape, where the magnet is located. In addition, two further quench zones are visible at either side of the magnetic defect, at the edges of the tape. This could very well be caused by the CFD architecture, where after the initial quench at the magnetic defect the current was diverted to a lower resistivity region around the edges of the tape and caused rapid heating. The sample was measured after the initial quench and no damage to the superconducting layer was discovered. This would seem to indicate that the prepared sample is on track to meet the requirements set forth by the FASTGRID project, aiming at developing HTS conductors that can withstand electric fields beyond 100 V m−1 without damage [107]. Further measurements on closely identical samples are currently ongoing at collaborating institutes to measure the NZPV electrically as well.

Figure 4.15: Sample 8 pulsed with a 650 A, 12 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a).

#### CFD architecture at Polytechnique Montréal

The measurement results of sample 5 are shown in Figure 4.16. It was prepared at Polytechnique Montréal with a CFD architecture [18] using a conductor similar to sample 4 as a basis and sent to KIT for NZPV measurement and for comparing results. Having an Ic of 85 A, this tape was pulsed with 85 A for a duration of 8 ms, where an NZPV of ∼2.2 m s−1 was measured at Polytechnique Montréal (see Table 4.1). Reproducing the same pulse at KIT using high-speed fluorescent thermal imaging did not yield the same results, however. While a clear electrical quench signal was visible as shown in Figure 4.16a, the thermal images did not capture a propagating quench zone. The only visible sign of the quench was the rapid development of a thin film boiling over the entire surface of the tape.

This behaviour is similar to that observed in metal stabilized tapes as previously discussed, where in the case of over-critical current pulses the current would divert into the electrical shunt layer and cause a rapid film boiling layer development. Here, however, there was no metallic shunt layer, only the standard 1.2 μm thick Ag coating together with the CFD architecture. Regardless of this, no reasonable NZPV value could be obtained for this tape with the thermal imaging method.

There is no clear explanation for the behaviour of this Ag tape. It is believed based on the existing experimental data, however, that the CFD architecture indeed forces the current to divert rapidly to the edges of the conductor, where it encounters a low resistivity path. Here it can flow again across both the length of the conductor and across the width of the superconductor face. This can result in a rapid surface boiling and in an almost instant, simultaneous surface boiling as opposed to a propagating zone, where the current can locally transfer from the superconducting layer into the silver layer directly above it. The results indicate that the CFD architecture does help distribute the heating over the tape's length and as a result it can be considered as a viable method for quench protection in HTS coated conductors.

Figure 4.16: Sample 5 (with a CFD architecture added at Polytechnique Montréal [18]) pulsed with a 85 A, 8 ms long pulse. (a) shows the applied current pulse and the recorded electric field across the tape. (b)-(d) show raw and thermal images corresponding to times indicated by the black vertical lines in (a). The red arrows indicate the quench development.

## 4.4 Summary

## 4.4.1 NZPV calculation validation

This chapter demonstrated that the normal zone propagation velocity can be obtained via the introduced thermal imaging method purely from the developed temperatures over the surface of an HTS tape. While in electrical NZPV measurements the quench propagation is calculated as the velocity of the developing electric field, in the shown thermal imaging approach the quench was defined as "the furthest point of the 93 K wave font". Consequently, the velocity of this wave front was calculated to determine the NZPV from a thermal perspective. Due to the development of film boiling regions and the inherent differences compared to the electrical NZPV measurement method, the shown thermal imaging cannot be used to calculate an NZPV in every case. The method is most suited in this respect to investigations until approximately 150 %Ic.

Measurement results on several HTS samples were presented with varying architecture in terms of substrate thickness, silver layer thickness as well as copper stabilization. Unfortunately data was not always available from both electrical and thermal imaging NZPV measurements and as a consequence the comparison is sparse. Nevertheless, several measurements indicate a reasonable match between the two NZPV measurement approaches, using different physical principles. For two Theva tapes, one stabilizer-free and another with a copper-nickel shunt both methods reported comparable NZPV values at closely identical current pulses. In another Theva tape measurement the sample quenched at a lower current level in the optical measurement and hence a lower NZPV was obtained. In general in the case of not matching results, the thermal imaging tends to under-estimate the NZPV reported by the electrical measurement method. Two such prominent exceptions were recorded as well, in stabilizer-free samples from SuperPower. These have shown a rather large discrepancy in the results between the two methods with a factor of 2 to 10 times slower NZPV measured by the thermal imaging. This was expected to a certain extent as the heating is a consequence of a developed electric field in the tape, and as such the temperature rise must lag a voltage signal. Simultaneously, the discrepancies between the two methods were observed in tapes with a critical current roughly half when compared to measurements where the two methods reported similar results. Since losses (and therefore heating) are proportional to the second power of the applied current, this may suggest that the thermally and electrically measured NZPVs are similar at higher current amplitudes due to the higher heating power during a quench. This also applies to a lower current amplitude, where a lower electric field may develop faster, followed by a slower heating due to a lower power dissipation.

Measuring the NZPV using fluorescent thermal imaging has its own limitations. Obtaining an NZPV for different current levels was difficult, especially for currents significantly above the Ic of the respective conductor. Using high current amplitudes the current rapidly transfers to the metallic shunt (if present) and causes quasiuniform heating over the entire observed surface, thereby preventing the formation of a propagating normal zone to be observed. In stabilizer-free tapes, on the other hand, a thin film boiling appeared over the entire surface at such high current levels, which then distorted the view and prevented the calculation of an NZPV. While the method is capable of reporting the normal zone propagation velocity in HTS tapes, it was not aimed primarily for this investigation. The method furthermore requires considerable manual post-processing that cannot be completely automated. As a result, while the thermal imaging method can serve as a comparison for NZPV measurement, it is not intended to replace other methods.

## 4.4.2 Effect of metallic stabilization

In Section 4.3.2 samples 3 and 9 with the 100 μm Cu-Ni and the 20 μm pure copper shunt, respectively, were discussed. It was shown that although the thickness of the metallic stabilization, as well as the composition, were different, the tapes quenched in a similar way with respect to the applied current pulse amplitude and duration. This is not to say that the NZPV or the stability of the tapes is independent of the metallic shunt thickness, rather that the qualitative appearance of quenched zones was observed in both tapes in a similar fashion.

## 4.4.3 Electric field development

When comparing the electrical measurement results of the Oxolutia sample with the CFD architecture from Figure 4.15a (sample 8) with that of e.g the SuperPower sample shown in Figure 4.8 (sample 4), the difference in the voltage profiles is immediately obvious. While the SuperPower sample developed the voltage gradually over the duration of the pulse, the Oxolutia sample answered with a step-like response to the applied current pulse. At the time of writing this was the first successful measurement on such a tape architecture and as such there was no data available from other institutes to confirm the reproducibility of the quench behaviour. The most likely explanation is related to the presence of the high inter-facial resistance between the superconducting and the covering silver layer. With a good electrical connection between these layers (in conventional HTS tapes), as soon as the superconductor's Jc is approached locally, a voltage can be recorded across the tape. Using a highly resistive CFD layer presents a barrier to current transfer into the silver layer, essentially forcing the current to stay in the superconductor. This could then cause Jc(T) to drop rapidly inside the superconductor, due to a tiny, dissipative zone and consequently cause a high electric field in accordance with the power law (Chapter 2). At a high enough voltage then current sharing between the silver and superconducting layer will take place and cause rapid local heating. In the shown measurement the electric field became so high that the current source could not maintain the required voltage to supply the requested 115 A.

Furthermore, on the example of the same SuperPower sample, shown in Figure 4.8, one can see that a relatively high electric field of up to 9Vm−1 was recorded, however no quench was visible on the thermal imaging. Simultaneously, at this "not quenched" current pulse of 110 A an energy of 1.98 J was absorbed by the tape. Then in the measurement where the tape did quench at 115 A, merely 0.57 J was absorbed over the first ∼7 ms of the pulse (until the appearance of the quench) and the electric field was 6.3 V m−1. This proves the existing difficulty that measuring an electric field over a longer distance does not necessarily give accurate information about the development of a localized dissipative zone. In general, a quench could be identified on the thermal images at an approximate electric field of 5Vm−1, however, this was highly dependent on the distance between the voltage taps.

On occasion, HTS tapes did burn out during measurement if the deposited energy was too high or concentrated in a single point. In most situations a burnout could be easily identified on the electrical signal, without inspecting the tape. In such cases the resistance of the tape section measured in between the voltage taps became too large, which resulted in a large, sharp voltage drop. As the maximum voltage of the current supply is limited, the current would consequently fall to a sustainable level. Such a measurement signal, especially with stabilizer-free tapes, indicated that the sample (most likely) suffered permanent damage. Tapes with sufficient metallic stabilization on the other hand may survive such a scenario due to their increased heat capacity. It was therefore surprising to observe that in the case of the Oxolutia sample (number 8) with the inkjet-printed CFD – despite the lack of a metallic stabilization – the sample did not suffer any damage even though the developed electric field was high enough to cause a drop in the supplied current.

### 4.4.4 Types of quench development

As shown in previous measurements, a quench can behave in several different ways. Probably the biggest distinction that can be made is how the quench spreads: starting from a single point or appearing at several locations almost simultaneously. Unfortunately, it was not possible to measure the NZPV in the current experimental setup simultaneously using both the optical and electrical methods. However, since the NZPV of HTS is routinely measured at other institutes for a range of current levels (including over-Ic) most likely the voltage profile during such a quench with multiple nucleation points is indistinguishable from that of a "normal", propagating quench zone. From the thermal images and results presented here, at the same time, it is clear that in such scenarios the heat distribution happened vastly differently.

As a conclusion, one can say that it is not only the deposited energy in the tape that is important, but how or rather what type of current pulse was applied, and consequently the rate at which the energy was deposited (heating power). It is already known that the NZPV of HTS is higher at higher current levels with respect to Ic [16], however the qualitative formation of normal regions was not investigated in detail so far.

N. Nguyen et al. [115] previously discovered non-homogeneous "banding" boiling over HTS and also showed that currents far above the tape's Ic caused a more homogeneous quench. Conclusions similar to this were drawn in this work as well and these can be summarized as follows. At "low" powers, corresponding to low amplitude current pulses, HTS tapes developed a single quench zone and depending on the duration of the pulse this zone then propagated. At "intermediate" heating rates, ∼30 % to 50 % above the Ic, coated conductors seemed to develop several quench zones almost simultaneously that rapidly expanded and caused heating and boiling of the coolant over the tape's (almost entire) surface. Then at even higher heating power and pulse amplitudes, the current transferred almost instantly into the metallic shunt and caused Joule heating over the entire tape surface.

This is an interesting phenomenon worth further experimental investigation. By designing HTS tapes with shunt layers having optimal heat capacity and resistivity characteristics, they can be fine-tuned for specific applications (e.g. superconducting fault current limiters). By adjusting these parameters, several simultaneous quench zones could be caused at the same time and if the temperature rise is kept in control, the tape could be held in the nucleate boiling regime with optimal heat transfer conditions (refer to Appendix, Figure A.3). In the current understanding, a quench nucleates in an HTS at a single point and spreads with a given NZPV along the length. This is supported by the standard electric NZPV measurements. In light of the appearance of multiple quench zones presented in this work as well, the quench and NZPV may not be defined as straightforward as previously thought. Although the NZPV of HTS tapes is undoubtedly not fast enough, by the appearance of multiple quench nucleation points, the energy could still be distributed and a thermal runaway prevented.

# 5 Fluorescent thermalimaging of HTS pancake coils

Previous chapters described the working principles of the fluorescent thermal imaging method when combined with high-speed recordings and also showed its application in quench investigation and NZPV measurement of single HTS tapes. This chapter shows the application of the method on the example of two distinct pancake coils wound from 2G HTS tapes: a conventional insulated and a non-insulated coil. The chapter details the design and preparation of the coils as well as the experimental setup, sets the aims for their measurement and gives a comparison of their final performance and behaviour in operation.

## 5.1 Experimental objectives

The stability and quench aspects of HTS coils were introduced previously in Chapter 2. As a proof of concept, fluorescent thermal imaging was implemented for measuring two HTS pancake coils, wound with different winding techniques. The first coil was a conventional insulated coil, whereas the second coil was wound without any turn-to-turn insulation.

Both coils have their own benefits and drawbacks and the aim of the investigation was to observe the coils during operation from a thermal and electrical perspective. Such a thermal imaging on insulated or NI coils was not carried out in the past and was not commonly available using state-of-the-art measurement methods. With the thermal imaging possible weak spots can be detected in the coil, coming either from design, material defects or simply due to exceeding the critical current of the coil locally due to some disturbance.

## 5.2 Electro-magnetic model

A number of numerical models exist using finite element modelling (FEM) and finite difference modelling (FDM) as well as equivalent circuit models [73, 86, 116–120] for simulating the behaviour and calculating the critical current of superconducting pancake coils including NI coils.

The development and simulation of a complete coil was beyond the scope of this work, however, FEM was used as initial analysis to determine the expected critical current, magnetic field as well as inductance and stored energy in the coils. A 2D axisymmetric, magneto-static model of the coil was created in the commercial, finite element method COMSOL Multiphysics 5.4 software, using the modified load line method [121]. It was implemented using COMSOL's magnetic field module, which uses the magnetic vector potential **A** as a state variable and the governing equations are

$$\mathbf{J} = \nabla \times \mathbf{H} \tag{5.1}$$

$$\mathbf{B} = \nabla \times \mathbf{A} \tag{5.2}$$

$$
\mathbf{J} = \sigma \mathbf{E}.\tag{5.3}
$$

In the model the HTS windings were simulated as homogenized superconducting domains of concentric rings as shown in Figure 5.1.

Figure 5.1: (a) shows a cross-section view of the coil turns and the air domain around it, where r is axis of symmetry. (b) shows a zoomed view of the of the coil geometry as highlighted by the red rectangle in (a), where each thin strip represents a turn of the pancake coil.

The FEM model takes the electrical conductivity σsc as input parameter, which was defined using the power law from e.q. 2.3. Furthermore, to avoid numerical instabilities – such as infinite or zero electrical conductivity – the final electrical conductivity of the superconducting domains was defined as

$$\sigma\_{\rm sc} = \begin{cases} \frac{J\_{\rm c}(B,\Theta)}{|E|} \left( \frac{|E|}{E\_{\rm c}} \right)^{\frac{1}{n(B,\Theta)}} + \sigma\_{\rm norm}, & \text{if } \sigma\_{\rm sc} < \sigma\_{\rm max} \\\\ \sigma\_{\rm max}, & \text{otherwise.} \end{cases} \tag{5.4}$$

Here Ec = 1 × 104 V m−1, σnorm = 1 × 106 S m−1 is the normal state conductivity and σmax = 1 × 1016 S m−1 is an artificial maximum limit for the conductivity of the superconductor. For simplification, adiabatic conditions were assumed at a temperature of 77 K. To simulate the magnetic field distribution around the coil an air domain was used as an approximation to the pure nitrogen environment in the planned experimental assembly.

The superconducting domain's critical current was modelled using as input the measured Ic(B, Θ) characteristic, as shown in Figure 5.2 [122]. The n-value was also available for various background fields and angles and was consequently applied in the model as n(B, Θ) (see Figure A.4 [122]).

Figure 5.2: Dependence of the critical current on the orientation and amplitude of the applied magnetic field in a the 4 mm-wide SCS 4050AP HTS tape used in the coils at 77 K [122, 123].

#### 5.2.1 Non-insulated coil

In the case of the NI coil, the current density may vary among the turns as current can bypass between them. Ignoring the resistive heating effects of such a bypass, a solution is sought where

$$J(x,y) = J\_c(x,y)(B,\Theta). \tag{5.5}$$

The critical current of the coil was then calculated by integrating the current density over all individual turns and finding the minimum

$$I\_{\rm c, \, min} = \min \int\_{n\_{\rm t}} J\_{\rm c}(B, \Theta) \mathrm{d}A\_{\rm t}, \tag{5.6}$$

and mean

$$I\_{\rm c, mean} = \text{mean} \int\_{n\_{\rm t}} J\_{\rm c}(B, \Theta) \text{d}A\_{\rm t},\tag{5.7}$$

values. Here nt represents a given turn number and At is the cross-section area of a tape. When applying Jc(B, Θ) as driving current to the NI coil, the minimum value of the surface integral current was 52.7 A whereas the mean value was 73.1 A. These are 93 % and 109 % of the experimentally obtained critical current of 57 A and 67 A (using the 0.1 μV cm−1 and 1 μV cm−1 criterions), respectively (see Table 5.1).

#### 5.2.2 Insulated coil

In the insulated coil additional boundary conditions were necessary to restrict the overall current in each turn to equality. In the first approach, the conductor was homogenized with the additional thickness of the Kapton® insulation1, resulting in a total thickness of 210 μm. In the second approach, both the HTS tape and the insulation were modelled, with the electrical conductivity of the insulator layer set to zero. To restrict current flow between turns in the homogenized approach

$$\mathbf{n} \cdot \mathbf{J}|\_{\Omega} = 0,\tag{5.8}$$

the boundaries were insulated electrically

$$
\sigma|\_{\Omega} = 0,\tag{5.9}
$$

1 DuPont™and Kapton® are trademarks or registered trademarks of E.I. du Pont de Nemours and Company.

where Ω denotes the boundary of the superconducting domains.

To find the critical current of the insulated coil with the given boundary conditions, a parametric sweep simulation was carried out, by applying uniform current from 0 A to 80 A to each turn of the coil. At each step the minimal critical current in the coil was calculated according to

$$I\_{\rm coil} = \int\_{n\_{\rm t}} J\_{\rm c}(B, \Theta) \mathrm{d}A\_{\rm t}. \tag{5.10}$$

By plotting the calculated critical current as a function of the applied current, the critical current of the coil was obtained at the point that satisfies e.q 5.5.

Figure 5.3 shows the results of the simulation searching for the coil's critical current in steps of 1 A. The critical current line represents the limiting critical current of the coil, when supplied with a corresponding applied current, Ic(B(Iapp), Θ). The intersection of the critical current and applied current line indicates the critical current of the coil as 74.5 A, representing a 3.5 % overshoot compared to the measured valued of 72 A (using the 1 μV cm−1 criterion).

Figure 5.3: Critical current search results using FEM simulation in the insulated coil. The curve shows the critical current of the coil as a function of the magnetic field, generated by the applied current as shown by the dashed line. The intersection of the two curves indicates the critical current of the coil.

## 5.2.3 Coil parameters

The magnetic field distribution around the cross-section of the coils is shown in Figure 5.4. To compare this with measurement results, the calculated magnetic field strength at position x, y = (0, 0.004) (location of the Arepoc Hall-effect sensor in the experiments) is plotted in the critical current measurements of the coils in Figure 5.15 and Figure 5.30 in Section 5.4 and Section 5.5. The calculated field strength of 250 mT to 260 mT at this point matches measurement results at currents corresponding to the simulated data, within error limits obtained in the Ic.

The stored energy of the coil ecoil in the FEM model was calculated as the integral of the magnetic field

$$e\_{\rm coil} = \frac{1}{2} \iiint \mathbf{B} \cdot \mathbf{H} \mathrm{d}v.\tag{5.11}$$

From this then the coil inductance was obtained using

$$L = \frac{2e\_{\text{coil}}}{I^2} \tag{5.12}$$

at the critical current of the coil, as calculated from the model.

A summary of the calculated values together with design parameters of the coil are presented in Table 5.1.

Figure 5.4: 2D magnetic field distribution over the cross-section of (a) the NI-coil and (b) the insulated coil. The y-axes represent the axis of symmetry.

Table 5.1: Specifications and properties of the NI and insulated coils, wound from SuperPower SCS4050 AP tape. † At 77 K, self-field. a Using the 0.1 μV cm−1 criterion. b Using the 1 μV cm−1 criterion. c Calculated value. ‡ Represents the "ideal" field coefficient calculated from the maximum magnetic field and current.

## 5.3 Coil preparation

This section discusses the main considerations in designing and building the experiment and winding the coils, focusing on the aspects of:

## 5.3.1 Cryostat

To avoid complex and expensive vacuum and cryocooler setups with optical windows – and since the aim was to test the coil only at 77 K – an in-house, 2-component cryostat was constructed. An aluminium container was placed inside a polystyrene container and the space between the two was filled up with liquid nitrogen, as shown in Figure 5.5. The coil was then placed in the aluminium box and a good thermal contact was made between the two with the help of a thin layer of thermal paste. This allowed the coil to be conduction-cooled to 77 K via the aluminium container's walls. The remainder space in the containers was filled with the evaporated, cold nitrogen gas, providing a cold and clear atmosphere, ideal for the optical measurement. The high-speed camera was then positioned above the setup and lowered close to the coil inside a separate optical cryostat that kept the camera at room temperature. The UV excitation light was also placed in the camera's cryostat, facing downwards. The setup was finally surrounded by multi-layer insulation from the top to prevent radiation heat loads, air movement and mixing with the cold nitrogen gas as well as to block out any ambient light that would interfere with the optical measurement.

The result was a conduction-cooled coil in a pure, cold, nitrogen atmosphere, where the continuous boil-off of liquid nitrogen provided a flow of cold gas keeping the atmospheric moisture out of the assembly. Cooling the coil from ambient temperature to 77 K was possible within 30 min as shown in Figure 5.6. However, to achieve complete thermal equilibrium in both the entire assembly as well as in the measurement devices, the coil was cooled for up to 3 h before the measurements started.

Figure 5.5: Cross-section drawing of the coil measurement assembly. The coil is fixed with thermal paste at the bottom of the aluminium container and excited by the UV LEDs from inside a separate cryostat housing the high-speed camera. Radiation heat load and ambient light are blocked by sheets of multi-layer insulation.

Figure 5.6: Temperature profile of the two current leads and the coil former when cooling the coil in the conduction-cooled assembly to 77 K. The dashed line indicates the liquid nitrogen temperature at atmospheric pressure.

## 5.3.2 Coil design

The coil design was governed by the need of a large, optically visible area during the measurement to be thermally imaged. Hence the inner coil former diameter was made as small as possible while the outer diameter as large as possible, keeping in mind the length of superconducting tape required and minimum bending diameter of the HTS tape [124]. As previously seen in Chapter 4 the boiling of liquid nitrogen severely impacts the recording quality and therefore it had to be avoided. As a consequence, the coil had to be conduction-cooled and the current leads had to be designed in a way not to obstruct the view of the coil.

Considering these aspects, a copper cylinder with a diameter of 30 mm and a height of 6 mm was prepared and served as the first current lead or contact. The cylinder was equipped with a slit from the outside until its centre line, where a 4 mm deep, M6 drill hole was present. To fix the HTS tape to the cylinder, the coated conductor was inserted into the slit in the cylinder and pushed into the centre hole. Here an M6 bolt with a vertical slit was pushed down onto the tape, which was then fixed by turning the bolt. This pulled in a portion of the tape and kept it mechanically stable for the first few turns of the winding. The bottom part of the same copper former had a 12 mm wide and ∼0.5 mm deep groove where a 12 mm HTS tape was soldered to provide the first current lead. This is shown in Figure 5.7. The outer diameter of the coil was fixed at 85 mm, which required about 30 m of HTS tape and was constrained by availability.

At the bottom and around the copper former, a "c-shaped", 2 mm thick copper disk provided mechanical and cooling support for the coil. This copper disk had a 30 mm hole

Figure 5.7: The back-side of the coil showing the 12 mm-wide HTS current lead as well as the coil's innermost windings soldered to the coil former.

at the centre – matching the diameter of the copper cylinder – as well as a 12 mm-wide slit in one direction to allow space for the HTS current lead. In the final assembly, the copper disk is insulated from the HTS current lead and the windings of the coil. The last two turns of the coil were soldered together, purely for mechanical reasons. The last turn of the winding was left to be ∼20 cm longer to be used as the second current lead.

The current contacts were two copper blocks with a multi-contact plug attached to them using an M8 thread and insulated from the aluminium container via a sheet of Kapton® foil. These were clamping the HTS leads about 20 cm away from the coil using a thin sheet of indium between the HTS surface and the copper block. In the experiments it was found that by thermally anchoring the current leads in the liquid nitrogen bath, the temperature of the current contacts did not rise above 78.5 K when applying current up to 74 A.

The insulated coil was designed according to the same specification and principles as the NI coil, however, due to the additional thickness of the electrical insulation, the coil parameters for an insulated coil could not be kept the same as those of the NI coil. It was decided to keep the number of turns constant. As a consequence, the inductance of the coil as well as the peak magnetic field and limiting critical current are slightly different. Due to the added thickness of the Kapton® tape, the coil outer diameter was 97 mm compared to 85 mm for the NI coil.

The winding was done exactly the same as in the case of the NI coil, with the only difference that the turns were insulated from one another by a 4 mm-wide and 50 μmthick Kapton® sheet. The instrumentation of the coil was done using the same voltage taps, Hall-sensor, the same current leads (including the Pt100 temperature sensors) and a new Pt100 temperature sensor for the centre of the coil, since this was bonded to the copper cylinder using X60, two-component adhesive.

## 5.3.3 Coil winding

For both coils, SuperPower, SCS4050 AP tape was used with a width of 4.1 mm and an effective thickness of about 163 μm, as governed by availability at our lab. Although the specifications of the tape state 40 μm copper stabilization on either side, the effective thickness of the copper layer is considerably thicker at some positions, as shown in Figure 5.8. Based on the dimensions of 30 mm inner diameter and 85 mm outer diameter, this resulted in a total of 157 windings.

The coil was wound using the copper cylinder as a former (or target) and a spool of the HTS tape on a source spool. Both were placed in level with one another to keep a consistent winding profile. To further keep the tape in position, a cylindrical nylon roll with a 4 mm groove around its circumference was used to guide the tape and smooth out any up/down swings resulting from the uneven alignment of the tape on the source spool. Pictures of the winding assembly illustrating the winding of the insulated coil are shown in Figure 5.9 and Figure 5.10.

Both the target coil and the source spool had a dc motor, the first for controlling the winding speed and direction and the latter for keeping tension in the tape by applying a constant torque. The dc stepper motor of the target coil was speed controlled by a LabVIEW software and the motor of the source spool was keeping tension in the tape, proportional to the applied current. The tape was aligned during the winding such that it was continuously, yet barely, touching the bottom copper disk and the result was a coil with a perfectly flat surface and a bottom flushed to the copper disk. This is shown in Figure 5.10a and the finished coil with an even surface in Figure 5.10b. Such an even surface was highly desired for an easier coating application.

Figure 5.8: Cross-section microscopy image of a SuperPower SCS4050 AP tape used for winding the coils [125].

Figure 5.9: Picture of the winding machine assembly, highlighting: (1) the source spool of the 4 mm HTS tape, (2) the spool of Kapton® being co-wound with the HTS tape and (3) the target coil. Both tapes were passed through a nylon cylinder (4) to keep them aligned during winding.

(a) (b)

Figure 5.10: (a) A coil being wound onto the 30 mm former, showing the c-shaped cooling plate as well. (b) The finished coil on the winding setup with the soldered outer joint. (c) A close-up picture of the Kapton® co-winding with the HTS tape. (d) A small material defect on one of the insulated coil's outer windings. The point defect is later also visible in the finished coil.

Jo et al. reported the effects of winding tension and correlated Ic reduction on layer-wound coils [126]. They found that coils wound with 4.9 N to 78.4 N maintained constant ∼90 % of the short sample Ic of the conductor. However, they went as far as 107.8 N, where considerable Ic reduction was already present. Although literature also suggests that higher winding tensions are possible, Kim et al. used 1 kg [74] and Yanagisawa et al. [127] 0.5 kg for winding a layer-wound coil and Jeon et al. [79] used 40 N. Initial experiments with 4 mm-wide pure copper tape indicated that higher tensions tend to strain the tape either on the top or along the bottom edge during winding. As a consequence, considering winding tensions reported in the literature and the need to avoid any possibility of damage, a relatively low winding tension of 2 kg to 2.5 kg was used.

## 5.3.4 Electrical joints

To provide good electrical contact and also to avoid high temperatures, a low melting point In-Ag (97 %-3 %) solder was used for all of the electrical joints. The melting temperature of this solder is 123 ◦C, which is below the temperatures that would permanently damage HTS coated conductors [128]. The soldering in all cases was done by first applying Spirflux 330® to the copper surface of the HTS tape or the copper former.

First, the copper cylinder was heated up and pre-tinned over the surfaces where the 12 mm HTS and the 4 mm HTS winding would later join. The 12 mm HTS current lead was also pre-tinned and soldered to the copper cylinder using a soldering iron. At this point the copper former – together with the soldered HTS current lead – was fixed on the winding setup. The 4 mm HTS tape was attached to the centre of the former and was pre-tinned over the length of the first winding. A 3 mm-wide In-Ag soldering tape was placed in between the copper former and the HTS coated conductor and the first few turns were wound to apply tension, similarly as shown in Figure 5.10a. The soldering at this point was done using a hot air blower until the soldering tape had visibly melted between the two surfaces.

The outer two turns were soldered in a similar way and the result is partially visible in Figure 5.9. Both the outer surface of the second last turn and the inner and outer surfaces of the last turn were pre-tinned. Again a 3 mm-wide In-Ag soldering strip was used between the turns, however in this case instead of a hot air blower a soldering iron was used. Going around the coil in the direction of the winding, the tape was heated until the solder melted between the outer two turns. This method allowed a good mechanical contact between the two last turns and avoided heating up the entire coil for soldering only a small section. Note that this also means that in the insulated coil, the last two turns are practically non-insulated as well.

The joint resistance was measured at current levels of 40 A and 50 A for both the insulated and non-insulated coils using a nanovoltmeter and was found to be approximately 200 nΩ and 600 nΩ, respectively. At these current levels the superconducting transition was not yet observable, but the higher current allowed measuring a voltage above the noise limit.

## 5.3.5 Coil instrumentation

A picture of the finished NI coil is shown in Figure 5.11 where the temperature sensors, voltage taps and the magnetic field sensor are shown. During measurements three temperatures were recorded: the two current leads, each with a Pt100 temperature sensor embedded in the copper blocks, and the central copper former. This latter had a 1.5 mm × 1.5 mm × 20 mm groove that allowed inserting a Pt100 temperature sensor. The sensors in the current leads were thermally bonded using silicone paste and fixed to the copper blocks by zip-ties. The sensor in the coil former was fixed in the groove using fast-drying, methyl methacrylate-based X60 glue. This served as both mechanical support as well as thermal bonding. For measuring the axial magnetic flux density Bz an Arepoc hall-effect sensor, with a 99.8 mV T−1 sensitivity, was fixed with double-sided duct tape at the centre of the copper cylinder, just above the M6 drill hole.

Figure 5.11: HTS non-insulated coil placed in the aluminium box and insulated from the bottom via Kapton® sheet. Three Pt100 temperature sensors, a pair of voltage taps, a Hall-effect magnetic field sensor as well as the area covered by the fluorescent coating are indicated by the white arrows. The black, dashed rectangle indicates the boundaries of the area recorded by the camera, in the same orientation as presented later in thermal images.

To avoid any distortion and additional unknowns in the coil, only a single pair of voltage taps were used, outside of the coil. These were soldered with In-Ag solder to the HTS current leads a few cm away from the coil. As a consequence, the voltage signal measured across the coil includes two resistive joints. First, between the 12 mm HTS current lead and the copper cylinder, and second, between the copper cylinder and the 4 mm windings of the coil. However, this resistance was measured to be insignificant for any practical purpose in this work.

The measurement devices used were closely identical to that described in Chapter 4 with the difference that for the coil several additional data acquisition channels were used, including a parallel-connected nanovoltmeter measuring the voltage drop across the coil. The Dewetron amplifier was used for four channels: current across an electrical shunt resistor, voltage drop across the coil, camera trigger signal reference and the magnetic field from the Hall sensor. In addition, three Dataforth SCM5B40/41 amplifier modules were used to acquire data from the Pt100 temperature sensors. All channels, except for the nanovoltmeter, were acquired using the NI USB 6289 DAQ.

In the case of the NI coil it was found that the Heizinger current source, used also for the pulsed HTS tape measurements, was accurate enough to carry out the experiments, despite the presence of some current ripples. These were mostly "filtered" by the RL characteristics of the coil. In the case of the insulated coil, however, even the smallest fluctuations in current caused large inductive voltage responses and hence made accurate measurements impossible. Therefore a more accurate, Cryogenic SMS120C, magnet ramping source was used that was specifically designed for inductive loads. The available ramping rates were slightly different for this power supply, with the maximum being 12 A s−1.

The insulated coil – after measured for various ramping rates – was retrofitted with a 5 Ω resistor (at 77 K) to be used as a local heater and test the coil's stability, shown in Figure 5.12. The same X60 glue was used as for the Pt100 temperature sensor at the centre of the coil and it was positioned approximately 1 cm away from the coil former, opposite to the side of the HTS current lead. The heater was deliberately positioned in the "body" of the coil, to allow the observation of heat flux, not necessarily at the centre of the coil.

## 5.3.6 Fluorescent coating

To observe local heating, the coil's surface was covered with the EuTFC-based fluorescent coating. Several tries were made including an airbrush and using a traditional paint brush to apply the coating to the coil's surface, however the result in all cases was not satisfactory as the coating became either too thin or highly non-uniform. For example,

Figure 5.12: A 5 Ω resistor retrofitted and used as a heater on the surface of the insulated coil.

using an airbrush proved to be difficult twofold: first, the coating has no colour and therefore it was difficult to see which part was coated already and how uniformly. Second, applying a successive layer over a coated area to improve the coating thickness is not feasible, as the acetone from the coating would dissolve any already existing coating. The only solution seemed to be the use of a single layer coating rapidly sprayed over the entire surface. However, due to the highly diluted coating solution, this method was not applicable and yielded poor results.

It was found that the same droplet coating method, as used when measuring single HTS tapes as described in Chapter 4 [129], worked best for the HTS coils as well and provided a sufficiently thick and uniform coating for the measurements. Figure 5.13 shows the quality of the final coating, where bright parts indicate either a higher optical reflectivity (e.g. copper surface) or a thicker fluorescent coating, whereas darker parts indicate a thinner fluorescent layer. Most of these optical artefacts were cancelled out during post-processing.

The initial concern with this method – applied to a pancake coil – was the possibility of the acetone-PMMA solution entering the space between the windings and forming an insulation at certain positions. However, in lack of a better approach for applying the coating and given the tight coil winding it was still adopted and no related problems were encountered later on. Nevertheless the coating was significantly more difficult than the coating of single HTS tapes. In the case of the NI coil, the coating did remain mostly on the surface, even if some of it entered in between the turns. With the insulated coil, however, problems were encountered due to the slight height difference between the HTS tape and the Kapton® co-winding. These two have effectively created a spiral-shaped groove on the surface of the coil. When applying the liquid coating these grooves hindered the distribution and caused highly non-uniform light intensity over

(a) Insulated coil (b) Non-insulated coil

Figure 5.13: Raw fluorescent image of (a) the insulated coil and (b) the NI coil. The image of the insulated coil already shows a retrofit resistor used as a heater.

the surface as shown in Figure 5.14. Increasing the amount of coating applied to the surface did not help as described above on the example of the air brush and at the same time caused more coating to disappear in between the turns. For a satisfactory coating it was found that instead of applying precise droplets continuously over the coil (inherently resulting in non-uniform coating) applying the coating from a larger height of 10 cm to 20 cm was more appropriate. This allowed the droplets to splash in a larger radius and self-distribute better over the surface while avoiding excessive coating solution build-up.

The high-speed camera, despite used at several thousand fps to measure quench in single HTS tapes (refer to Chapter 4, [129]), was used at the slowest speed setting of 50 fps, which was found sufficient for measuring the heating in the coils. Depending on the position of the coil and the camera, this allowed an approximate recording time of about 10 min to 12 min per experiment. In the post-processing stage the recordings

Figure 5.14: (a) shows the NI coil's surface completely covered with the coating solution. In (b) the fluorescent emission from the coating shows a strong non-uniformity and partially missing spots.

were generally averaged over 200 ms to reduce noise, without any loss of information about the transient effects.

## 5.4 Non-insulated coil measurement results

#### 5.4.1 Critical current

The coil was initially measured in liquid nitrogen to safely determine the operating margins. Figure 5.15 shows the results of this Ic measurement, where using the 0.1 μV cm−1 and 1 μV cm−1 criterion the critical currents were determined as 57 A and 67 A, respectively. The magnetic field development, as simulated by the FEM model, shows a lower expected field than experimentally measured. This is due to the difference in current densities in the model and experiment as shown on the calculated Ic values in Table 5.1. After concluding all experiments presented in this section, another Ic measurement was performed to check whether the coil was damaged in any way. It was found that the critical current of the coil did not degrade in any measurable quantity.

#### 5.4.2 Turn-to-turn resistance

The importance of the turn-to-turn contact resistivity in NI coils was introduced previously in Chapter 2. By knowing the winding tension, Rc can be calculated from eq. 2.17 provided that measured data is available for Rct. In this case one must assume

Figure 5.15: Initial Ic measurement of the NI coil showing the 0.1 μV cm−1 and 1 μV cm−1 electric field criterion. The critical current of the coil is then 57 A and 67 A, respectively. The plot shows data points extracted from the measurement and not the actual ramping plot.

that the initial measurements are accurate enough as well as that the winding tension is constant during winding. For measuring the turn-to-turn contact resistivity, the same setup was used as by Otten [130]. Two pieces of HTS tape were placed face-to-back on top of one another with an overlap of 4 cm, and were put in a hydraulic press and cooled to 77 K in liquid nitrogen. The hydraulic press was operated manually by applying increasing pressure on the tape and, simultaneously, a 10 A dc current was applied across the overlap of the two tape pieces. The pressure dependence of the resistance was obtained by measuring the voltage drop across the interface between the two HTS tapes as well as the pressure by the use of strain gauges. To accurately reproduce the condition of the demonstrator coils, the samples were taken directly from the spool and were not cleaned before measurement.

The measurement results in Figure 5.16 show a surface contact resistivity of around 300 μΩ cm2, which lies in the range reported by Otten [130] in a more extensive measurement using the same face-to-back configuration. When substituting this value into eq. 2.17 to calculate the coil resistance numerically

$$R\_{\rm c} = \sum\_{i=1}^{156} \frac{3 \times 10^{-8} \,\Omega \,\text{cm}^2}{2 \cdot \pi \cdot (0.0015 \,\text{m} + i \cdot 163 \times 10^{-6} \,\text{m}) \cdot 0.004 \,\text{m}} = 6.6 \,\text{m}\Omega. \tag{5.13}$$

However, deriving the characteristic coil resistance using eq. 2.9 with the discharge time constant and inductance of the coil resulted in

$$R\_{\rm c} = \frac{1.38 \times 10^{-3} \,\mathrm{H}}{610 \times 10^{-3} \,\mathrm{s}} = 2.23 \,\mathrm{m}\Omega. \tag{5.14}$$

Figure 5.16: Surface contact resistivity measurement at 77 K on two pieces of SCS4050 AP tape from SuperPower during a single load cycle.

The two equations show a three-fold difference in the coil resistance when calculated using the two different approaches. This suggests either different surface characteristics between the measured tape sample and the long length HTS usedin the coil,measurement inaccuracies or some fundamentally different behaviour.

## 5.4.3 Thermal and electrical effects of various ramping speeds

In the conduction-cooled experiments, the coil was ramped using several ramping rates, while being recorded at 50 fps to observe whether the different ramping rates caused any heating. The tested rates were 0.1 A s−1, 0.5 A s−1, 1As−1, 2As−1, 5As−1, 10 A s−1, 20 A s−1, 50 A s−1 and 100 A s−1 with the target current levels being 40 A, 50 A and 55 A. A representative part of the current and magnetic field developments are shown in Figure 5.17 during coil ramping.

The electrical results of a single, typical ramping measurement are shown in more detail in Figure 5.18 where the coil was ramped with 10 A s−1 to 55 A2. In each measurement the coil was ramped with a specific speed and upon reaching the target current, it was given 60 s to reach equilibrium. This was then followed by a ramp down, with the same speed, followed again by a 60 s final relaxation time. In order to allow the coil to return into thermal equilibrium after each measurement, a ∼5 min beak was kept between any two experiments and the liquid nitrogen level was continuously monitored.

Figure 5.17: Time development of current and magnetic field in the non-insulated coil during the first 60 s when ramped at various ramping rates to 55 A.

2 For simpler visual representation, the inductive voltages from the measurements are always represented as positive in the figures.

Figure 5.18: Time development of measured current, voltage and magnetic field in the non-insulated coil when ramped at 10 A s−1 to 55 A. The radial and the spiral current (Ir, Is) were calculated using eq. 5.15.

Due to the inductance and lack of turn-to-turn insulation of NI coils, during a current ramp with an arbitrary dI/dt a voltage is induced according to eq. 2.11. While normally this would act against current flowing in the spiral path, in an NI coil the current may also bypass the turns across the turn-to-turn resistances. Hence the higher the inductance and the ramping rate, the more current will initially bypass the turns – or the coil itself – and revert back into the spiral path with a delay. The inverse happens in a sudden discharge, where the current supply is suddenly disconnected from the coil, e.g. by means of a switch (Figure 2.10), and as a consequence the coil sees a large dI/dt and develops a large voltage. This voltage then forces currents to flow through both the superconducting spiral path as well as the turn-to-turn resistances and dissipate quickly.

Assuming that it is predominantly the current in the spiral path Is that generates the magnetic field, the current in the radial path Ir may be approximated as

$$I\_{\mathbf{r}} = I\_{\text{coil}} - \frac{B}{k}.\tag{5.15}$$

Here k is the coil field constant defined as

$$k = \frac{B(t)}{I(t)}.\tag{5.16}$$

Since for an NI coil the magnetic field has a delay, k was obtained experimentally after the magnetic field has stabilized as

$$k\_{\parallel} = \frac{B\_{\text{max}}}{I\_{\text{max}}} = 0.003 \,\text{98 T A}^{-1}.\tag{5.17}$$

Knowing the amount of current bypassing through the turn-to-turn resistances, the heating P and absorbed energy Q may also be calculated using

$$\begin{split} Q\_{\text{coil}} &= \int\_{0}^{\infty} P \, \text{d}t \\ &= \int\_{0}^{\infty} I\_{\text{r}}^{2} R\_{\text{c}} \, \text{d}t \\ &= \int\_{0}^{\infty} \left( I\_{\text{coil}} - \frac{B}{k} \right)^{2} R\_{\text{c}} \, \text{d}t. \end{split} \tag{5.18}$$

This equation, however, may only apply to the ramping and steady-state operation, when the coil is still in its superconducting state. After the superconductor transition begins, the energy stored in the magnetic field of the coil will also begin dissipating and it is more appropriate to use

$$\begin{split} Q\_{\text{coil}} &= \int\_{0}^{\infty} P \, \text{d}t \\ &= \int\_{0}^{\infty} I\_{\text{coil}} V\_{\text{coil}} \, \text{d}t. \end{split} \tag{5.19}$$

However, calculating the heat load using eq. 5.18 may result in uncertainties due to the uncertainty in the value of Rc itself as pointed out in Section 5.4.2. While several groups have proposed complex equivalent circuit grid models to simulate heating in NI coils [86, 117], the development of such models is beyond the aims set forth in this experimental investigation. Instead, to approximate the current sharing between the turns, the charging profiles were compared with the coil field constant and the peak current bypass value was calculated from eq. 5.15. The results are shown in Figure 5.19 for all ramping measurements. The plot shows that – as expected – a faster ramp causes higher (peak) radial bypass currents, with less than 5 A of bypass until a charging rate of 2 A s−1.

In three of the (time-wise consecutive) measurements the coil's lower-right sections showed signs of localized heating, later developing into a semi-circle and disappearing during ramping down. In these three experiments the coil was ramped with 50 A s−1, 100 A s−1 and 0.1 A s−1 to 40 A, 40 A and 50 A, respectively. The measurements are shown in Figure 5.20-Figure 5.22 with selected time excerpts of the thermal imaging along with the electrical measurements. As visible, most of the coil remained at close to

Figure 5.19: Estimated current bypass across the turn-to-turn contact resistances, calculated using eq. 5.15, at various current ramping speeds.

77 K, with the exception of a select area. In Figure 5.20b and Figure 5.20c the spreading of the warmer zone is visible just after the end of the ramp and approximately half way through the 60 s plateau. This warmer zone then dissipated when the coil discharging began. Similar effects are visible in Figure 5.21b and Figure 5.21c showing the 100 A s−1 ramping measurement, where heating is still visible around the same zone as in the previous measurement, although to a lower extent.

The effects are more expressed in the measurement ramping to 50 A as shown in the thermal images Figure 5.22b and Figure 5.22c. Here, apart from the distorted ring-shaped heating in the coil, it would seem that the heating is concentrated to the path directly between the inner and outer current contact, despite the slowest ramping rate tested.

Generally, in some of the measurements with target currents of 40 A and 50 A the coil's innermost windings showed moderate heating (80 K to 85 K) around the centre copper former. However, the same heating was not observed when targeting 55 A. In this latter case, some overall heating was observed in the coil body, yet to very moderate levels of ∼80 K and without any deductible patterns with regards to ramping speed.

As a summary, in select measurements different parts of the coil were observed to warm up to different levels, however no clear pattern was recognized with respect to ramping rate or target current level (below the coil's Ic). Therefore, from the thermal imaging results, it cannot be clearly stated whether faster ramping rates would incur in increased heat load in an NI coil. However, it has to be kept in mind that the investigated coil's inductance and characteristic resistance are low and may be the cause of this result. While approximately a 5 min recovery time was kept between all ramping measurements to allow the coil returning into a steady state, it is possible

Figure 5.20: Electrical and thermal effects of 50 A s−1 ramping rate to 40 A.

that this time window was not sufficient, which may also induced anomalies into the measurement.

Figure 5.21: Electrical and thermal effects of 100 A s−1 ramping rate to 40 A.

Figure 5.22: Electrical and thermal effects of 0.1 A s−1 ramping rate to 50 A.

## 5.4.4 Charging delay and charging time constant

Figure 5.23 shows the summary of all experimentally obtained charging delays of the coil when ramped to 40 A, 50 A and 55 A with ramping rates ranging from 0.1 A s−1 to 100 A s−1. The trend of the charging delays is clearly visible, and as expected [75], a faster ramping rate resulted in a longer time delay. The 0.1 A s−1 measurement shows negative values, however, this is merely an artefact from the calculation. At such a slow ramping rate the field increases closely with the coil field coefficient and hence at the time when the ramping ends, the 99.3 % field criterion is already reached. During data analysis this then may result in negative values. Interesting is also that with higher target current, the charging delay was moderately, yet consistently shorter. This can be caused by the field saturation of the coil, where more current simply does no longer cause a linear rise in the field and subsequently causes a faster perceived settling time. The plot also shows a saturation effect of the charging delay, where the difference between 10 A s−1 and 100 A s−1 charging only causes a moderate few second difference.

In Figure 5.24 it is shown that the coil charging time constant (as introduced in Chapter 2, [75]) is in fact a function of the ramping speed as well, where the higher the ramping rate, the shorter the time constant. This is supported by the fact that a faster ramping rate causes a higher voltage according to eq. 2.11, which in turn causes more current to flow in the radial direction, through a larger number of turn-to-turn contact resistances. This effectively increases the coil characteristic resistance and hence results in a lower time constant in eq. 2.9. Interestingly, Kim et al. [75] observed a correlation between the charging time constant and the target current level. By using

Figure 5.23: Data for the 0.1 A s−1 ramping rate to 55 A is missing.

Figure 5.24: Charging time constant τup of the NI coil as a function of ramping rates between 0.1 A s−1 to 100 A s−1. Data for the 0.1 A s−1 ramping rate to 55 A is missing.

a constant ramping rate in a racetrack pancake coil, they found that ramping to higher current levels caused a longer time constant. This was not observed in this work.

## 5.4.5 Sudden discharge

The coil discharge time constant was previously introduced in Section 2.4.2. To determine this experimentally, as well as to determine if any local heating is present in an accidental discharge, the coil was ramped to 30 A and 40 A in two successive measurements, where it was held for 60 s to allow for the centre magnetic field to stabilize. After this, the switch shown in Figure 2.10 was opened, allowing the coil to rapidly dissipate the stored magnetic energy in the coil resistance. The measurement results in Figure 5.25 show that the discharge time does not depend on the current amplitude and it is approximately 610 ms. This observation matches that of Wang et al. [76], where the same independence of the discharge time constant from the current amplitude was observed. Additionally, so have Kim et al. [75] come to the same conclusion with current amplitudes of 20 A, 40 A, 60 A and 80 A.

The energy stored in a coil at a given current level was calculated as

$$e\_{\rm coil} = \frac{1}{2}LI^2.\tag{5.20}$$

By knowing the discharge time constant the power dissipation was then calculated as

$$P\_{\text{discharge}} = \frac{e\_{\text{coil}}}{\tau\_{\text{down}}}.\tag{5.21}$$

Figure 5.25: Normalized magnetic field development during sudden discharge tests at steady DC current levels of 30 A and 40 A. The dashed line indicates the 0.37 criterion where the time constant is evaluated.

From this the energy stored in the coil when held at 40 A was 1.1 J and the power dissipation approximately 1.8W for a duration of 610 ms. With a coil mass of ∼0.270 kg and by approximating the heat capacity with that of pure copper at 77 K as 195 J kg−1 K−1 [52, 131], the overall heat capacity of the coil would be 52.65 J K−1. Disregarding heat distribution and the effects of cooling, it is clear that the energy released during a sudden discharge was not enough to cause any heating in the coil. This is merely a consequence of the small size and inductance of the test coil. In a larger coil with significantly larger inductance local heating may occur and make additional protection measures necessary.

#### 5.4.6 Overcurrent

Measurements were done above the limiting critical current of the coil to test its behaviour without any artificial defect as to observe how a "natural" thermal runaway would happen. In overcurrent measurements, the coil was first ramped to a previously validated safe current level and from there slowly to higher currents. Every few amperes the ramping was stopped and the coil held steady to observe if the coil voltage was still stable. This stepwise ramping was then continued until the coil voltage would no longer stabilise and a "natural" runaway would start to appear. Special care was taken to manually ramp down the coil in case of a high voltage to permit future experimenting and avoid any damages. In different measurements, the transition point was observed at slightly different current levels, which suggests that slight disturbances in cooling, ramping and current levels can influence the thermal stability of the NI coil.

#### Overcurrent at 72 A

In one of the experiments the coil was initially ramped to 71 A at a rate of 10 A s−1 and the voltage was monitored, as shown in Figure 5.26. This eventually stabilized and so the coil was ramped higher to 72 A, this time at a slower 1As−1 rate, as shown at time 160 s. At this current level a slow, yet steady voltage shift was observed and so the coil was held at this level.

Since the magnetic field also reduced with the increasing voltage, the current must have been changing from the spiral to the radial path. This is also supported by the quasi-linear voltage rise, as expected with increasingly more current flowing through a resistor. In Figure 5.26b it is visible that initially the outer turns of the coil were at a higher temperature, presumably due to slightly worse cooling. At the same time, distinct heating is visible in the form of circles with varying width (Figure 5.26c and Figure 5.26d, as shown by the white arrows), indicating that over some turns the current was bypassing across the turn-to-turn contacts and caused losses. However, at about 298 s into the measurement, the voltage dropped and the magnetic field increased sharply, indicating that the current reverted — for an unknown reason – back into the spiral path quite abruptly. Simultaneously, the centre windings of the coil heated up, as shown in Figure 5.26e, where the ring-like hot areas were replaced by a uniformly warmer zone, mostly on the left-hand side of the coil. This showed that current flowing though the turn-to-turn contacts of the coil can be seen on the thermal image as heating over possibly several turns, which is most likely a function of the cooling performance and the distribution of local defects in Jc along the length of the wound conductor.

## Overcurrent at 74 A

To reproduce previous results and observe the same behaviour again, the coil was ramped in a subsequent measurement directly to 72 A at a rate of 10 A s−1 as shown in Figure 5.27. However, in this case no voltage drift was observed. The coil was therefore ramped to 73 A and finally 74 A at a slower 1As−1 rate, where the coil voltage became unstable and started increasing once again.

Over the course of two minutes at steady 74 A, the coil voltage was rising continuously up until its final value of 0.47 V, where the coil was manually ramped down quickly to prevent any damage to the superconductor windings. During the transition, the magnetic field started decaying with the developing voltage, indicating a slow current transfer into the radial path. Then after 337 s a much sharper voltage rise and accompanied field drop is visible, likely due to (some of) the HTS tapes becoming resistive.

In Figure 5.27b it is visible that the outside of the coil is again at a higher temperature, however still cryostable. Warmer zones are still visible around the centre and middle layer of the coil, indicating heating due to currents bypassing across the turn-to-turn contacts. Figure 5.27c shows that the middle layer seems to have cooled back mostly and at the same time, due to the voltage rise of the coil, the innermost windings started heating up. At the bottom of the image, where the outer current lead arrives at the coil, a hot spot is also slightly visible. Then in Figure 5.27d distinct heating is visible around the centre coil former. Finally in Figure 5.27e, just before the coil was ramped down to protect it, the hottest temperature of around 100 K was reached at the few innermost windings of the coil. Note that the heating is limited to the innermost perhaps 1 cm of windings and that the outer turns are virtually unaffected. It is also interesting to observe that at the time when the coil voltage started rising, the middle turns of the coil actually cooled down and when the voltage continued to rise, the central windings started heating up instead. Most likely when the coil voltage is low, only the "major" defects in the coil would be bypassed, however at higher voltages more current could be forced across all turn-to-turn contacts, thereby reducing the current in the superconductor as well as local heating. Due to the larger overall thermal mass this could then manifest in a form of a local cooling effect. At the same time, the centre windings see the combined effect of superconducting to resistive transition of the HTS due to the reduced temperature margin and high magnetic field, combined with the current bypass and Joule heating due to the resistive joints.

(a)

(d) t = 298.80 s (e) t = 300.0 s

Figure 5.26: Electrical and thermal measurement results of the 72 A overcurrent measurement. Dashed lines indicate positions of thermal images. (a) shows the time development of the applied current ramp, coil voltage and magnetic field. (b)-(e) show thermal images, where white arrows indicate specific features.

Figure 5.27: Electrical and thermal measurement results of the 74 A overcurrent measurement. Dashed lines indicate positions of thermal images. (a) shows the time development of the applied current ramp, coil voltage and magnetic field. (b)-(e) show thermal images, where white arrows indicate specific features.

## 5.4.7 AC current ripples

Most dc power supplies have a certain current ripple on the output either coming from the grid frequency or due to the rectification of the current. In rotating machines the changing magnetic field induces ac current in conductors that have in principle a similar effect. Such ripples often appear at higher multiples of the power line frequency and are commonly 1 % to 3 % of the nominal current. For these reasons the effect of ac current ripples was also investigated on the test coil using various frequencies and ripple amplitudes, generated according to eq. 5.22.

$$i(t) = I + I\_{\text{ripple}} \cdot \sin(2\pi ft) \tag{5.22}$$

where I was a steady dc current between 30 A and 50 A, Iripple was between 1 A and 10 A and the frequencies, f, ranged from 0.1 Hz to 400 Hz. The measurements were recorded both via fluorescent thermal imaging as well as electrically, making sure that the data acquisition speed was fast enough to capture high frequency signals with satisfactory resolution. By monitoring the change of magnetic field at the centre of the coil simultaneously to currents, the actual amount of current bypassing the coil – and as such flowing through the radial path – was determined using the coil field coefficient. Table 5.2 shows the measured voltage and magnetic field peak-to-peak fluctuations with a 1 A and 10 A ac current superimposed on top of a 30 A dc current in the coil with frequencies from 0.5 Hz to 400 Hz. Figure 5.28 then shows the fluctuations of the magnetic field with only the 10 A ac signal.

The variations in field are too weak in the case of the 1 A ac signal, however they are clearly visible with the 10 A ac current waveform. With the coil constant of 3.98 mT A−1, a 20 A peak-to-peak signal – if completely following the spiral path inside the coil – would generate magnetic field fluctuations of

$$k \cdot i\_{\text{max}} = 0.0796 \,\text{T} \tag{5.23}$$

in the measurement. Any lower change in field then indicates that the current is proportionally bypassing the spiral path and flows through the turn-to-turn contact resistances to some extent. The results in Table 5.2 indicate that this is the case in most higher frequencies. Even at 1 Hz the field change is merely 32 mT, representing about 53 % of the change, in case if the current were to completely follow the spiral path. However, this is not unexpected, since even this frequency already represents an effective ramping rate of 40 A s−1.

Further experiments using the same principle of dc offset current, as high as 50 A (∼80% Ic), with modulated ac signals up to 100 Hz in frequency and 1 A, 2 A and 10 A

Table 5.2: Peak-to-peak variations in central magnetic field and induced voltage with a 1 A and 10 A ac signal modulated on top of a 30 A dc current, for frequencies from 0.5 Hz to 400 Hz. † Shows the percentage of the ac current part bypassing the spiral path completely as calculated from the coil field coefficient.

in amplitude are shown in Figure 5.29. The coil was ramped with 5As−1 to the target current, where the ac current was applied at 1 A with frequencies from 0.1 Hz to 100 Hz. After this the ac signal amplitude was increased and the signals repeated. At the end the coil was ramped down using the same ramping rate as during charging.

It was found that the 1 A ac current had no effect on the coil, however at 2 A during the 25 Hz ac current, a visible hot zone was detected in the coil as shown in Figure 5.29b and Figure 5.29c. The heating around the coil's centre (indicated by the white arrows) was already present after ramping the coil. A hot zone on the upper right side of the coil appeared during the 25 Hz signal. This was shortly followed by 50 Hz and 100 Hz ac currents and as such the warmer region of the coil did not recover and a temperature offset was visible until the end of the signals. After this, the warmer region started to recover. No heating was observed during ac signals with 10 A in amplitude for frequencies between 0.1 Hz to 5 Hz and the coil was slowly recovering until the ramp down as shown in Figure 5.29d.

At the exact time when the coil discharging began at the end of the experiment, a brief yet distinctly visible heating was detected once again in the same zone of the coil as before, shown in Figure 5.29e.

Due to the limited thermal imaging recording length, the 10 A 25 Hz, 50 Hz and 100 Hz measurements were done in a subsequent measurement. While this measurement was

Figure 5.28: Normalized magnetic field during an alternating current with an amplitude of 10 A applied on top of a steady 30 A dc current. (a) shows frequencies between 0.5 Hz to 25 Hz and (b) frequencies between 50 Hz to 400 Hz. In (b) the power supply experienced initially a small drop in the dc offset current, which was manually corrected for in this plot.

merely a few minutes after the previous one, no warmer zones were observed with only minor heating during the experiment. A sudden heating at the beginning of the ramp down was also not observed in this measurement.

Figure 5.29: Electrical and thermal measurement results of 1 A, 2 A and 10 A ac current from 0.1 Hz to 100 Hz on top of a steady 50 A dc current. (a) shows the electrical measurements with a normalized magnetic field. Arrows indicate the amplitude of the applied ac current. (b)-(c) show initial heating during the 2 A ac signal between frequencies 25 Hz to 100 Hz. (d)-(e) show the coil partially recovered and sudden, local heating at the moment of ramping down.

## 5.5 Insulated coil measurement results

## 5.5.1 Critical current

The insulated coil was initially also measured in liquid nitrogen to obtain the critical current of the coil. The results are shown in Figure 5.30, with the critical current determined as 65 A and 72 A using the 0.1 μV cm−1 and 1 μV cm−1 criterion, respectively. Due to the added thickness of the Kapton® insulation between the turns, the magnetic field generated by the coil is reduced and hence the critical current is marginally higher, compared to the NI coil. After measuring the coil's critical current in liquid nitrogen and confirming the characteristic properties, further measurements were carried out in the conduction-cooled experimental assembly.

## 5.5.2 Thermal and electrical effects of various ramping speeds

The coil was ramped with 0.1 A s−1, 0.5 A s−1, 1As−1, 2As−1, 5As−1 and 10 A s−1 ramping rates until 40 A, 50 A and 60 A while both thermal images as well as the magnetic field and coil voltage development were recorded. Unfortunately, despite the more accurate magnet ramping current source, due to the inductance of the coil and the minor current fluctuations, the noise in the fast DAQ system was still too high to do reliable measurements. Hence, instead of the DAQ signal, the nanovoltmeter was used at a higher speed with the readout range set to 1 V for sufficient accuracy.

Figure 5.30: Initial Ic measurement of the insulated coil showing the 0.1 μV cm−1 and 1 μV cm−1 electric field criterion. The limiting critical current of the coil is then 65 A and 72 A, respectively. The plot shows data points extracted from the measurement and not the actual ramping plot.

The inductance of the coil was determined using eq. 2.11 and averaging over several ramping measurements was determined as 1.44 mH.

Figure 5.31 shows the magnetic field development of the coil with varying current ramping rates. Compared to the NI coil (refer to Figure 5.17) it is visible that this coil has no charging delay, the magnetic field is directly proportional to the applied current.

Due to the zero resistance path of the superconducting windings and the lack of turn-to-turn resistive bypass paths, no distinct heating was observed during the ramping measurements. The used power supply did not allow the generation of ac waveforms and the use of extended overcurrent was also avoided due to accidental burnout. Instead, a different approach was used to try and determine the stability of this coil.

#### 5.5.3 Heat pulses

As the coil voltage rose relatively quickly around currentlevels of70 A,going considerably above the critical current was avoided in ramping measurements. Instead, to detect any anomalies in the coil and observe its behaviour in transient states, a heater was added on the surface and short heat pulses were applied. This was previously introduced in Section 5.3.5 and shown in Figure 5.12.

The heat pulses were carried out using varying heating power, duration and at different current levels. The three current levels used in the coil were 40 A, 50 A and 60 A and the heater was pulsed with approximately 1.25W and 3.2W for durations between 1 s to 7 s. In the measurements the coil was ramped using a rate of 5As−1, where upon

Figure 5.31: Time development of current and magnetic field in the insulated coil when ramped with a rate between 0.1 A s−1 to 10 A s−1 until 60 A.

reaching the target current a 60 s stabilization period was allowed. Thereafter, the heater was pulsed with the selected power for durations incremented by 1 s at a time. After each heater pulse the coil was ramped down using the same 5As−1 ramping rate.

### Heating before detectable voltage

In the measurements up to 60 A coil current and heat pulses of up to 2 s at a power of 3.2W no voltage was detected between the coil's terminals. Figure 5.32 shows the results of this experiment where the heater was pulsed for a duration of 2 s with a nominal current of 800 mA. As visible in Figure 5.32a there was no detectable voltage signal in the coil. However, looking at Figure 5.32b and Figure 5.32c, a minor amount of heat transfer is already visible from the heater into the coil. While this is indeed barely visible in the static image, in the time development of the post-processed measurement the signal can be identified. This indicates that while heat disturbances were visible, these do not necessarily cause magnetic field fluctuations or any voltage response, provided that the local critical current density of the coil is not exceeded by the transport current.

#### Normal transition

In the heater pulse measurements the coil remained stable until heater current pulses of 800 mA (3.2W) for a duration of 6 s, with a total heat dissipation of 19.2 J in the resistor. While a clear voltage drop was measured, the coil recovered after each heat pulse ended. A normal transition was first observed when the pulse duration was increased to 7 s, providing about 3.2W of heating power and an overall energy of 21.4 J. This measurement is shown in Figure 5.33a. The heater was started just before 81 s, which is visible in Figure 5.33b as the heating of the resistor changes the reflected light and hence causes a false thermal reading, nevertheless, useful for timing the heater. Approximately 2.5 s after the heater was started, a voltage signal was also detected. However, after turning the heater off, the coil did not recover and a thermal and electrical runaway began. This is visible in Figure 5.33c, where the heater was already off, yet the temperature kept increasing. Figure 5.33d then shows that the heating is concentrated predominantly to the turns touched by the heater and that propagation to neighbouring turns is negligible. In this image one can see a hot ring reaching temperatures of 150 K. At this point the sharp transitional voltage caused an overflow in the nanovoltmeter and simultaneously caused the power supply to trigger quench protection and drop the current to zero. As such, the coil was protected and the heat started to diffuse further along the turn as well as into neighbouring windings as shown in Figure 5.33e.

Figure 5.32: Electrical and thermal measurement results of a 2.5 s long heater pulse in the insulated coil. (a) shows the time development of the applied current ramp, coil voltage, magnetic field as well as the heater pulse. (b) and (c) show thermal images, where white arrows indicate specific features.

Unlike in the case of the NI coil, where the normal transition was slow and accompanied by a detectable reduction of the centre magnetic field, in this insulated coil the field remained unchanged even with increasing voltage. A lack of electrical bypass paths led to large voltages developed and quasi-local heating over the involved windings, which – if undetected – may have destructive consequences.

After the quench, the coil's critical current was re-measured and was found to be unchanged.

Figure 5.33: Electrical and thermal measurement results of the 7 s long heater pulse in the insulated coil. Dashed lines indicate positions of thermal images. (a) shows the time development of the applied current ramp, coil voltage, magnetic field as well as the heater pulse. (b)-(e) show thermal images, where white arrows highlight specific features. Colour changes of the heater are optical artefacts and should be ignored.

## 5.6 Summary

Two pancake coils were wound from HTS coated conductors into an approximately 3 cm inner and 10 cm outer diameter coil. The first coil was wound without turn-to-turn insulation whereas the second coil used Kapton® as an insulating layer between the turns. The surface of both coils was coated with the fluorescent EuTFC-based coating with the aim to observe local heating in the conduction-cooled assembly during various operating phases.

## 5.6.1 Non-insulated coil

Electrical measurements on the NI coil have shown than when ramping with current rates between 0.1 A s−1 to 100 A s−1 to 40 A, 50 A and 55 A (60 %, 78 % and 85 % of the coil's Ic respectively) the coil charging delay is a function of the ramping rate, where a faster ramping rate increases the charging delay. This delay, however, became increasingly saturated above 5As−1. At this charging rate the coil needed approximately 16 s to reach 99.3 % of the target magnetic field from the moment where the ramping has ended. From sudden discharge measurements it was also determined that the discharge time constant of the coil is 610 ms and it is not dependent on the current amplitude from where the coil was discharged.

Fluorescent thermal imaging measurements have not revealed a consistent increase in heating at increased ramping rates. While heating was present in some parts of the coil during some of the ramps, due to the lack of repeatability these may be called anomalies related to the cooling performance inside the experimental assembly or some other undetermined factor. Similar results were observed in the case of the sudden discharge measurements, where no reasonable heating was detected. Due to the size and stored insignificant amount of energy in the coil, no heating was present in a sudden discharge. It is believed that larger coils with considerably higher inductance would behave differently in many scenarios when compared to the small test coil prepared in this work.

Overcurrent measurements of the coil revealed an anomalous behaviour, where a continuously increasing coil voltage suddenly recovered. In the measurement the magnetic field started dropping, accompanied by an increase in coil voltage, both signalling an increasing current transfer into the radial path. However, instead of a catastrophic thermal runaway, the coil current redistributed and both voltage and magnetic field levels returned to normal. The effects were visible on the thermal imaging, which revealed rapid current redistribution (and heating redistribution) at the moment of sudden recovery.

In yet another overcurrent measurement the coil was ramped to 74 A (110 % of the critical current) and kept there until the beginning of a thermal runaway. The thermal imaging measurements clearly showed that the origin of the normal zone was at the centre of the coil, where a resistive joint exists between coil former and superconductor. Simultaneously, this is the point of highest magnetic field. Over the duration of several seconds, only the few innermost turns heated up to approximately 100 K. The coil was manually protected by ramping it down and measuring its Ic after the quench revealed no degradation. The measurement has shown that the thermal runaway was virtually constrained to the location where it has started with little to no propagation in the radial direction.

Applied ac current to the coil, modulated on top of a steady dc current, mostly bypassed the spiral path and hence caused only minor fluctuations in the magnetic field. Heating due to the ac current was also not visible at ac amplitudes of 1 A and 10 A on a 30 A dc current due to the low characteristic coil resistance. A 10 A current ripple (which is already significant 15 % of the coil's Ic) would cause merely 0.2W of heating when approximated as I2R losses with a characteristic coil resistance of 2.23 mΩ, also assuming a complete and uniform bypass of the spiral path. Sudden heating as a result of a 2 A 25 Hz to 100 Hz ac current on top of a 50 A dc current became visible in one of the measurements. In the same measurement, the coil has shown local heating in the same position when ramping the coil down. Interestingly, however, this behaviour could not be reproduced in a subsequent measurement with the same frequencies using a higher, 10 A ac signal. The local heat generation suggests that ac current bypass may not happen completely uniformly in a coil, instead it could be concentrated to parts of the coil with slightly different turn-to-turn resistance characteristics.

## 5.6.2 Insulated coil

In the prepared insulated coil, no distinct or characteristic heating was observed when ramped from 0.1 A s−1 to 10 A s−1 until different current levels. To avoid accidentally burning out the coil in an overcurrent measurement, instead a resistive heater was mounted on the coil surface with the aim to observe the stability and the normal transition behaviour. The coil withstood a continuous heat input of 3.2W for up to 6 s, where both the electrical and thermal signs of instability were observable. Nevertheless the coil recovered on its own after the heater was turned off. When the heater pulse duration was increased to 7 s, a thermal runaway began. Starting from the heater's position a hot zone developed, which spread predominantly along the turns in contact with the heater in a semi-circle shape. Neighbouring turns were mostly unaffected by the thermal runaway. The coil was automatically protected by the current supply's built-in quench protection mechanism. Measuring the coil's critical current after the thermal runaway did not reveal any damage in this case either.

# 6 Conclusion and outlook

In this work a 2D temperature mapping method was developed for quench analysis in HTS tapes and applications. The application of high-speed fluorescent thermal imaging was demonstrated at 77 K on several HTS tapes as well as two HTS pancake coils, one of which was wound as a non-insulated coil.

Various HTS tapes were measured with pulsed currents to observe quench development and propagation in liquid nitrogen caused by a magnetic defect. The conductors varied in manufacturer, stabilization as well as architecture, including the novel current flow diverter concept. By defining the normal zone from a temperature perspective as areas above the critical temperature of HTS, a normal zone propagation velocity could also be calculated from the thermal imaging method. Cross-validating the measurement results with data provided on identical tapes by Polytechnique Montréal, Canada revealed that at higher operating currents the NZPVs calculated from electrical and thermal readings match more accurately than at lower currents. In the latter case the thermal imaging consistently reported slower NZPV, with up to one order of magnitude difference. This mismatch between the NZPVs was ascribed to the quadratic dependence of heating power on the applied current, P = I2 · Rsc. While a fast electric field development was possible in HTS tapes with low critical currents, the low heating power in these tapes caused the thermal zone propagation to lag the electrical signal. On the other hand, in tapes with a higher critical current (and hence higher applied currents), the considerably higher heating power caused the propagation of the thermal zone to match closely that of the electric field.

The appearance of multiple quench zones was visible in several tapes at current pulses slightly above the critical current. The appearance of these additional quench zones followed the initial quench almost instantaneously and the behaviour is ascribed to the presence of natural defects. The existence of a quench in this manner may not be immediately obvious from electrical measurements and indicates that the concept of quench and normal zone propagation in HTS is not necessarily as straightforward as a single propagating zone at a given speed. The presence of additional quench zones have a big influence on the heating distribution in the tapes.

The test coils were measured using various current ramping rates and current levels. Due to a low coil resistance and critical current, even a fast ramping rate of 100 A s−1 did not cause any visible heating. However, this is most likely the consequence of a small demonstrator coil. In a larger coil, where the ramping would take considerably longer, resistive joints as well as currents bypassing through the turn-to-turn resistances would be expected to cause higher losses and possible heating. Investigating ac current ripples on top of a steady coil current in the coil yielded the same results, where the low currents involved made it difficult to observe any heating related to the ac current.

The NI coil was also investigated in overcurrent scenarios, where the coil was ramped until 72 A and 74 A (corresponding to 108 % and 111 % of the coil's critical current) and held steady to observe a thermal and electrical runaway. In the 72 A measurement a voltage shift was observed across the coil's terminals, which was accompanied by a steady drop of the central magnetic field over the duration of several seconds. This was a clear indication of an increasing fraction of the current bypassing the spiral path and flowing across the turn-to-turn resistances. At one point, however – without any intervention – the coil voltage fell back to previous levels and the magnetic field also increased in a step-like manner. On the thermal images a clear redistribution of the heating load was visible at the time of the sudden change. The exact reason for this behaviour is currently not clear and requires further investigation.

In a subsequent measurement the NI coil was ramped to the previously validated current level, however at this time the coil remained stable. An instability became only apparent when slowly ramped to 74 A, where the coil was held for a longer period. During this measurement the coil voltage increased at a slow, yet steady rate and eventually a runaway was recorded where the coil was quickly ramped down. On the thermal image clear heating was visible around the centre windings of the coil, which – over the course of a few seconds – did not spread in the axial direction in the body of the coil, but remained constrained to the innermost few windings. Temperatures up to 100 K were measured at this position. This behaviour was expected in an NI coil without any artificial defect, since the centre was the position of highest magnetic field coinciding with a resistive joint in the design.

The results indicate that while NI coils are highly stable, their behaviouris not completely understood yet. While no natural weak spots were detected during operation, in the case of an overcurrent the heating did occur locally and was not distributed rapidly.

Similar to the NI coil, the insulated coil was also thermally imaged in operation, where no reproducible heating was visible in the coil during various ramping speeds. Instead, to test the stability of the coil and the capabilities of the thermal imaging method, a resistive heater was mounted on the surface of the coil (retrofitted) and current pulses of varying length and amplitude were applied to the heater. With increasing heater pulses the coil was destabilized using a heater power of 3.2W for 7 s at a dc coil current of 60 A. Here a local hot spot development was observed after the heater pulse at the position of the heat input and a temperature of 150 K was measured before the coil was protected. The heating was concentrated to the initiation of quench and did not spread in the transverse direction. A temperature rise was visible along the turns in a half-circle shape, constrained to the windings initially disturbed by the heater.

While the wound demonstrators coils have shown no heating due to local defects or reasonable heating during various ramping measurements, this may be due to the involved low coil resistance and current levels. Larger coils with significantly higher inductance would incur a higher heat load during ramping and therefore are expected to behave differently.

The introducedmethod has the potential formeasuring temperatures in superconducting applications. Beyond the work presented in this thesis, several improvements are possible that would make the technique more applicable and useful for specific scenarios. While the current coating technique was suitable for flat tapes, coating uneven surfaces – even that of the coil – proved challenging. Researching another method could make the coating process simpler and more reliable, however, it would most likely require a novel solvent for the Europium fluorescent dye, which would require further investigation regarding performance at cryogenic temperatures and fluorescent properties. In this work, due to the required fast, transient effects to be recorded, a high-speed camera was used. The working principles of this camera prevented the method to be used in real-time. Simultaneously, due to the different target market of the used camera, it has incurred in significant noise at low light intensities. By the choice of the camera (for example higher digital resolution), the method could be made more appropriate for specific measurements where accuracy is of higher importance than acquisition speed. In this case, the method could also be implemented in true real-time using simultaneous post-processing.

# A Appendix

Figure A.1 shows the 3D-printed frame for the optical filter. Since such narrow bandpass filters are commonly made for optics applications with given standard sizes and mounting options, finding these in larger diameter for a camera lens front-mount proved difficult. Using a custom-made frame allowed using a standard optical filter in between the camera lens and the camera's CMOS sensor and enabled blocking a wide spectrum of unwanted light.

UV lights for microscopy measurements often use light beams that illuminate a small surface of only a few mm2. Commercial UV light fixtures used for curing, on the other hand, while precise in emission spectrum are often extremely expensive and bulky. As a compromise, an array of LEDs was used as shown in Figure A.2 built from an aluminium heat sink and several UV LEDs already mounted on a starboard PCB.

The heat transfer between a submerged object and boiling liquid nitrogen at 77 K is highly non-linear as shown in Figure A.3. The optimal heat transfer is between

Figure A.1: The 3D-printed frame allowing to insert the optical bandpass filter in between the camera and the lens.

Figure A.2: Array of LEDs on a starboard PCB used as excitation light in the experiments. The LEDs were connected in series and driven with their nominal current of 700 mA.

10 K to 20 K temperature difference. In practical applications and during a quench of superconductors, exceeding this limit severely limits the cooling performance due to the formation of a thin film boiling region that acts as an insulation.

Figure A.4 shows the dependence of the n-value on the applied magnetic field and angle [122]. This data was used in the commercial FEM software COMSOL Multiphysics to calculate the coils' critical current.

Figure A.3: Typical heat flux between liquid nitrogen at a temperature of 77.3 K and a submerged object [106].

Figure A.4: Dependence of the power-law n-value on the magnetic field angle and amplitude, used in simulations [122].

# List ofFigures

# List of Tables

# List of Symbols

#### Overview of HTS tapes and their stability

## Bibliography

[131] N. J. Simon, E. S. Drexler and R. P. Reed: Properties of copper and copper alloys at cryogenic temperature. Volume NIST monograph 177. National Bureau of Standards, 1992. Chapter 7, pages 1–15 (cited on page 109).

## **Karlsruher Schriftenreihe zur Supraleitung**

Karlsruher Institut für Technologie (KIT) | ISSN 1869-1765

## Band 032 Roland Gyuráki Fluorescent thermal imaging method for investigating transient effects in high-temperature superconductor tapes and coils. ISBN 978-3-7315-1064-2

## **Karlsruher Schriftenreihe zur Supraleitung**

Prof. Dr. Tabea Arndt, Prof. Dr. rer. nat. Bernhard Holzapfel, Prof. Dr. rer. nat. Sebastian Kempf, Prof. Dr.-Ing. Mathias Noe (Hrsg.)

A principal barrier to a wide-scale application of high-temperature superconductors (HTS), such as rare-earth-based coated conductors, is the difficulty of quench detection and protection. Quench refers to the abrupt transition of a superconductor from a dissipation-free state into a normal conducting regime where heating occurs due to the appearance of a finite electrical resistance. Such an event can damage a superconducting tape or device locally by causing a sudden temperature rise potentially reaching above the melting point of the constituting materials.

This work presents the development and application of highspeed fluorescent thermal imaging, a novel technique for quench detection and analysis. Using a fluorescent dye in the form of a surface coating – with a precisely measured, temperature-dependent light emission – the surface temperature changes in HTS can be calculated from variations in light intensity. The developed technique allows for the measurement of rapid transient events, such as quench propagation in HTS tapes, as well as the quantitative observation of localised hot spots over larger 2D surfaces with sub-millisecond temporal resolution.

The applied technique uncovered peculiar effects in novel HTS tape architectures and also helped to verify and better understand the hot spot development in both insulated and non-insulated, HTS-wound pancake coils when operated in various conditions at 77 K.

032

**R. Gyuráki Fluorescent thermal imaging of transient events in HTS tapes and coils**