21, rue d’Artois, F-75008 PARIS A2-201 CIGRE 2014 http

* [email protected]

Short Circuit Verification for a 570 MVA, 420 kV single-phase GSU-Transformer by SC-Withstand Tests on a Mock-up Unit

Gerald Leber1*, Helfried Passath1, Mohamed Ryadi2, Patrice Hurlet3

1Siemens AG Österreich - Transformers Weiz Research Elingasse 3, 8160 Weiz, Austria

2EDF - R&D

1 Av Général de Gaulle, 92141 Clamart, France 3 EDF- CNEPE

8 Rue de Boutteville, 37200 Tours, France

SUMMARY Large generator step-up transformers in power generation plants are exposed to several kinds of stresses under service conditions. One kind of stress is caused by forces at short circuits, which can even endanger and affect the entire reliability of a power plant. According to the international standards, there are three methods to demonstrate the ability to withstand the dynamic effects of a short circuit. The first one relies only on calculation and design checks with the manufacturer´s design rules for short-circuits. The second one observes the performance during a conventional full-scale short-circuit test and the last one involves a design comparison to a similar reference transformer which has successfully passed a short-circuit test. EDF and Siemens decided to evaluate the short-circuit capability of a 570 MVA single-phase GSU-transformer by comparison to a similar reference mock-up transformer designed, manufactured and short-circuit tested for this purpose. The reference transformer was designed to undergo equal or similar mechanical stresses as in the real full-scale unit during a short circuit. In this paper the most significant stress types in windings are analysed, described and evaluated. Similarity criteria and constraints regarding radial and axial winding stresses are discussed in detail. The short circuit tests on the mock-up transformer were performed in two test sequences. The first one was a successfully performed “customer acceptance test” where nominal short circuit currents and stresses were applied to demonstrate the short-circuit stresses in the real unit. The second test series was a “destructive test”. Currents shots were increased step by step until the unit experienced a notable change in the winding reactance. The purpose of the second test was to identify the critical mechanical stresses which caused permanent winding deformations and loss of stability on the mechanical structure. Results from a visual out–of tank active part inspection and some failure modes detected after winding dismantling are presented in this paper. Frequency Response Analysis (FRA) measurements performed before and after the short-circuit test series show high correlations to detected mechanical movements. A not well known failure mode of helical windings, the so called “spiralling effect”, where compressive forces tighten up the winding turns towards the inner support, are analysed. Corrective measures to prevent such effects in transformer windings are suggested.

The paper presents an experimental as well as a theoretical design evaluation and calculation approach to verify the short-circuit withstand capability of large power transformers unable to be tested in high-power laboratories.

KEYWORDS Transformer – Short circuit withstand test – Short circuit forces – Short circuit verification – Spiralling

21, rue d’Artois, F-75008 PARIS A2-201 CIGRE 2014 http : //www.cigre.org

1

1. Introduction During the last centuries power ratings and transmission voltages have increased and therefore dielectric, thermal and also mechanical aspects in transformers have become more pronounced. While dielectric and thermal tests can be performed on each individual transformer unit and the acceptance tests are well defined by the IEC and ANSI international standards, the short-circuit testing of large power transformers is still done only in very rare cases for new transformer designs. For very large GSU (Generator-Step-Up) transformer there are often limitations regarding high power test laboratory capabilities and constraints regarding transport, delivery time and costs. In most new transformer design cases it is sufficient for the utilities to trust the manufacturer´s long term short-circuit design and manufacturing experience and mostly only a short-circuit design review is conducted to reinforce the purchaser confidence in the vendor´s design and manufacturing capabilities. In the case of GSU transformers installed in large power plants, which are vital components in the power generation and transmission systems, any elementary risk of poor short-circuit performance during system fault conditions can cause a long term outage of the whole plant and can result in a huge financial loss. Therefore, in such high potential risk cases, it might be reasonable to find a method to demonstrate the short circuit performance of such units by means of a short circuit test, when a full-scale test is not possible. Recently, EDF and Siemens decided to demonstrate the withstand strength of a 570 MVA, 420 kV single-phase GSU-transformer by means of a short-circuit test on a mock-up transformer. The mock-up was designed and manufactured by the Siemens Weiz factory and short-circuit tested at KEMA in order to verify the short-circuit performance of the real 570 MVA unit, which could not be full-scale tested due to the high short-circuit power demand and short lead time restrictions. The method of demonstrating the ability to withstand short-circuits chosen by EDF and Siemens was based on the international IEC standard 60076-5 and Annex A and B [1] with some additional requirements and tests specified between the user of the equipment and the manufacturer. There are three alternative methods to demonstrate the ability to withstand the dynamic effects of a short circuit in the IEC standard. The first one relies only on a calculation and design check against the manufacturer´s design rules for short-circuit strength. The second one checks the performance during a conventional short circuit test on a full-scale unit to be evaluated. The last method compares the design to a similar reference transformer, which has successfully passed a short-circuit test. The latter method is always controversial among experts in this field. It is always heavily discussed to what extent mechanical similarity can be and needs to be demonstrated for a reference transformer unit. The IEC 60076-5 Annex A and B, an informative guideline, gives some guidance how a comparative evaluation with a reference transformer can be conducted. However, in practice, there often remain some uncertainties how to evaluate the short-circuit performance by a similar tested mock-up or full-scale unit. The following chapters give a practical approach on how such a comparative short-circuit design evaluation can be performed in a way that both vendor and purchaser are satisfied with the demonstration of the ability to withstand short circuits. 2. Definition of similar transformer unit/mock-up for short circuit verification The EDF technical specification CST 72.C.042.02, clause 5.1.6 [2] specifies how the short circuit withstand strength has to be checked for any new transformer design. The manufacturer has to describe the mechanical design of the apparatus in a short-circuit design file and it must also compare the new design to an apparatus that has successfully passed a short-circuit strength test. The similarity requirements are given by the following 5 items:

1. Same winding technology, for example layer, coil, disk or wafer 2. Same type of conductor, for example transposed epoxy resin impregnated conductor 3. Relative stresses (ratio between actual stress and critical stress) of ±20% of those apparatus to

be qualified 4. Axial pre-stress and peripheral torque value of ±20% of the apparatus to be qualified 5. Successful short-circuit strength test of the similar apparatus

2

Although these requirements are based on the IEC 60076-5 Annex A and B and are less strict in terms of similarity characteristics, discussions between purchaser and manufacturer concluded, that it would be impossible to find a similar short-circuit tested transformer that completely fulfilled the 5 items. Although the Siemens Weiz plant successfully tested a three phase 440 MVA, 400 kV GSU transformer in the past [3], it was not possible to use this unit as a similar reference unit because of conceptual differences in the winding arrangement (2 axial-split LV winding systems) and non comparable conductor types. Therefore both parties decided to design, manufacture and test a mock-up transformer that is able to fulfil the similarity requirements and to demonstrate short-circuit withstand capability of the real 570 MVA unit.

3. Design parameters of 570 MVA GSU and similar mock-up transformer The most important design characteristic of this 570 MVA single-phase transformer is that a so called multi-concentric winding arrangement was chosen. This was to fulfil the low short- circuit impedance and transport weight requirements by the customer. Besides this, the multi-concentric winding arrangement (HV1–2 layer LV–HV2), as shown in figure 1, also has some advantages such as lower radial compressive hoop stresses in the double-layer LV winding. This is due to reversing axial flux effects and additional support of the outer highly compressive stressed LVo winding by the inner tensile stressed LVi winding.

Fig.1: Winding arrangement for the 570 MVA unit and the mock-up transformer The inner HV1 winding was a helical layer winding with spacers, the LV winding a double layer helical winding with spacers and the HV2 winding was an axial split disk winding with tapped disks and voltage inlet at the axial winding centre. The two HV windings were connected in series. As listed in table 1, the mechanical design of the mock-up unit was made to have the same mechanical and geometrical parameters in terms of winding arrangement, core and winding diameters, winding types and conductors, number of conductors in the HV disks as well as the number of radial and axial support spacers at the windings. In principle the mock-up unit can be considered as a quarter winding length cut out from the real 570 MVA GSU-unit. In terms of winding turn ratios, axial windings lengths and radial spacer widths there are some differences in order to obtain similar mechanical stresses on the conductors during the short-circuit test.

3

Table 1: Electrical and mechanical parameters of mock-up and 570 MVA GSU-transformer Characteristic design parameters Mock-up transformer 570 MVA transformer Type Single-phase oil immersed

mock-up transformer Single-phase oil immersed power transformer

Rated Power 300 MVA (SC Power Rating) 570 MVA Rated voltage HV 37.6 kV 405/√3 kV ± 2.47% Rated voltage LV 3.23 kV (LV short circuited) 20 kV Frequency 50 Hz 50 Hz Cooling Type - ODAF SC impedance; Min/Nom/Max - 14.57/14.73/14.94 % SC impedance 4.7 Ohm - Core type 1-leg core with flux-return 4-leg core Core diameter 1262 mm 1262 mm Core window height 2570 mm 1170 mm Number of wound legs 1 2 Number of windings per leg 3 (HV1, LV, HV2) 3 (HV1, LV, HV2) Tap changer - Off-load Number of turns HV1/LV/HV2 20/11/89 158/46/365+2x14 HV1 i/o winding diameters 1380/1460 mm 1380/1460 mm LVi i/o winding diameters 1566/1698 mm 1566/1698 mm LVo i/o winding diameters 1730/1862 mm 1730/1862 mm HV2 i/o winding diameters 2064/2308 mm 2064/2308 mm Number axial sticks(supports)/spacers 60 60 Effective winding height HV1/LV/HV2 210/526/520 mm 1834/2086/2144 mm CTC for HV1

45 (3.75x1.52) mm; Rp0.2=170 N/mm²; Epoxy resin bonded


CTC for LVi and LVo

77 (6.47x1.15) mm ; Rp0.2 = 200N/mm²; Epoxy bonded

77 (6.47x1.15) mm; Rp0.2 = 200 N/mm²; Epoxy resin bonded

CTC A for HV2



CTC B for HV2



Twin conductor for HV2

2 (16.8x2.56) mm ; Rp0.2=170 N/mm²; Epoxy resin bonded


Some simplifications were made for the mock-up transformer by designing a simple flux return yoke frame instead of a real round shaped yoke. Also only one wound leg was made instead of the two wound legs for the four legged 570 MVA unit (figure 2). As the clamping and tie-bar concept was of the same type and the two electrically parallel connected winding legs are short-circuit stressed in the same way, both simplifications do not influence the mechanical and electrical characteristics during a short-circuit. Fig.2a: Active part of mock-up transformer Fig.2b: Active part of 570 MV GSU (one wound leg and flux- return frame yoke) (2 wound legs and flux-return legs)

4

The tapped section disks were also provided for the mock-up transformer but were not wired out with external connections because it was possible to get similar radial and axial stresses for the tapped disks, as explained in chapter 4. Lead exits and leads were also designed identically in the surroundings of the windings, where radial winding stray flux cause forces at the leads. 4. Comparison and evaluation of the most important stresses during short-circuit During a short circuit there are a huge number of mechanical stress types in a transformer. The most important ones have to be addressed when designing the mock-up unit to prove its short-circuit withstand capability. The most critical stresses during short-circuit in a transformer are:

• Radial compressive hoop stresses • Windings subjected to compressive hoop stresses may fail because of either inward over-

bending of conductors at the spans between two consecutive radial supports (forced buckling) or loss of concentric form stability resulting in radial in- and outward deformations of conductors at the winding circumference (free buckling). Also “spiralling” is caused by compressive hoop stresses in pitched helical or layer windings, which lead to a thrust force on the lead exit. Spiralling can occur in helical windings when an “open turn lead exit” at the winding ends is subjected to radial forces and the conductors tend to shift circumferentially by tightening up on the radial support underneath and by twisting itself towards a smaller diameter higher-pitch winding.

• Radial tensile hoop stresses • Tensile stresses occur in the conductor cross section when the winding is subjected to radial

outward forces. • Axial compressive stresses on spacers and winding conductors • Winding conductors and spacers are subjected to compression due to cumulated axial forces

at the winding centre. Those forces have their origin from axial winding-end forces generated by the radial stray flux at the winding-ends.

Although there are several stresses like radial and axial bending stresses in conductors between supports, stresses at end-stack insulation, clamping plates and tie bars, the stresses in table 2 are the most important to be considered in the design. Table 2: Comparison of maximum stresses in 570 MVA GSU to stresses in mock-up Tx

HV1 LVi LVo HV2 570 MVA

GSU Mock-up Tx

570 MVA GSU

Mock-up Tx

570 MVA GSU

Mock-up Tx

570 MVA GSU

Mock-up Tx

Compressive hoop stress [N/mm²] 54 58 - - 104 105 - -

Tensile hoop stress [N/mm²] - - 25 22 - - 119 119

Tangential trust force at lead exit (Spiralling) [kN]

11 13 -7 -3 347 374 - -

Axial compressive stress on spacer [N/mm²]

12 8 22 18 22 18 29 26

Axial clamping pre-stress on spacer [N/mm²]

4.5 4.5 4.5 4.5 4.5 4.5 4.5 4.5

Figure 3a and 3b gives the magnetic field plot at short circuit for the 570 MVA GSU and the mock-up transformer. The magnetic stray flux density in the main gap, between the low voltage outer-layer (LVo) and high-voltage winding 2 (HV2), is of the same magnitude in both transformers.

5

Fig.3a: 570 MVA GSU (max-tap) Fig.3b: Mock-up transformer Electromagnetic field plot at short circuit Electromagnetic field plot at short circuit For this reason the most critical stresses, the compressive hoop stress in the LVo layer and the tensile stress in the HV2, are the same. The compressive stress at the inner high-voltage winding (HV1) is about 10% higher and the tensile stress at the inner low-voltage winding (LVi) is about 10 % lower in the mock-up compared to the GSU. Due to the fact that the mock-up transformer is much shorter in the axial winding length, there are more radial and reversing flux components at the inner windings at a multi-concentric winding arrangement. These effects also have some influence on the axial cumulative forces in the windings and therefore the compressive axial stresses on the spacers are about 11% lower in HV2 and 18 % lower in the LV layers of the mock-up transformer. Although these axial stresses in the windings are lower compared to the stresses in the real GSU transformer the absolute value is quite low for epoxy bonded conductors and also far below the allowable compressive stress withstand limits on pressboard spacers. Conductor tilting is also not a matter of concern, as the strands of the CTC´s and twin conductor were glued together with epoxy resin. The axial compressive stress in the HV1 winding deviates by more than 20% in the mock-up transformer. For this reason, it was agreed between user and manufacturer to demonstrate the withstand capability of the GSU-transformer´s HV1 winding by an application of a static load at winding processing in the factory which is equal to the axial short-circuit force. In figure 4 the most critical radial compressive hoop stress in the LVo epoxy bonded CTC´s over the total winding height is given. In the real GSU transformer, the stresses for the extreme tap positions are slightly different in each individual conductor due to the tapped disk in the HV2 which influences the magnetic ampere turns over the winding height. The radial compressive stress in the mock-up unit was designed in a way to get the same maximum radial compressive hoop stress in the conductor as in the worst case short circuit situation of the real GSU-transformer. The tangential trust force at the upper and lower winding end, an important parameter to evaluate spiralling effects, is slightly higher in the mock-up unit compared to the force in the real GSU-transformer. Axial compressive stresses on winding conductors and spacers are within the ± 20% tolerance limits and are not so critical in this design, as already mentioned above.

6

Fig.4: Axial and radial stresses in conductors over winding height of LVo helical winding

Fig. 5: Axial and radial stresses in HV2 disks over total winding height Figure 5 shows the mean tensile stress in the HV2 winding disk over the total winding height for both units. Due to two different CTC conductors for the tapped and main sections and a twin conductor at the winding centre, the mean tensile stress in each individual disk is discontinuous over the winding height. Nevertheless, the mock-up unit was designed in a way that the three different conductor types have the same maximum tensile stress during a short circuit as the real 570 MVA GSU-transformer. Finally, it can be concluded, that the design of the mock-up transformer is representative to show similar stresses like in the real GSU-unit during short-circuit. The similarity requirements specified by EDF are met. 5. Short circuit test on mock-up transformer and visual inspection results In order to show the short-circuit withstand capability of the mock-up transformer it was decided to perform two different short-circuit test series. At first a customer acceptance test according to IEC was carried out to demonstrate that the unit is able to withstand a short circuit that can occur in the real GSU transformer during service. After the acceptance test the oil in the tank was drained and an in-tank visual inspection was made directly at KEMA site to check the condition of the leads, clampings and outer windings. Finally a destructive short circuit test was performed in order to demonstrate the safety margins of the design. A visual inspection and dismantling in the factory took place to study and analyse the final failure mode.

7

5.1 Acceptance short circuit test on mock-up transformer (1st test series) The short-circuit tests were performed using the pre-set test method. The supply voltage was applied on the HV winding after the LV winding had previously been short-circuited. The rated asymmetrical short circuit current is the 100% reference to obtain the stresses in the mock-up transformer as given in table 2. Table 3: Asymmetrical short circuit peak current and measured reactance change for 3 shots

Reactance change during short-circuit acceptance tests

Kema test number

ÎHV SC

X before test X after test Reactance change

measured [kA peak] [%] [Ω] [Ω] [%]

100421-5008 23 103.3 4.822 4.823 0.03 100421-5009 -22.3 100.1 4.825 0.08 100421-5010 22.9 102.8 4.831 0.20

According to IEC 60076-5 three shots are required for a single-phase transformer. The tolerance in the asymmetrical peak current is ±5% and the test circuit was adjusted in a way to get minimum100% and maximum105% test current. The tests in table 3 showed no remarkable increase in the reactance and also no abnormalities regarding fault current, Buchholz relay tripping or gas generation. The final reactance change was only +0.2% of +1% allowed according to IEC. The maximum test current was 103.3 % of the specified one and due to the squared relation between current and force, the stresses were at 107% of those given in table 2. 5.1.2 Visual in-tank inspection at high power laboratory site A visual in-tank inspection was performed in order to check the overall mechanical condition of the test unit. Although only the external mechanical condition of the active part can be checked, it is essential to inspect the leads and lead support structure. Mechanical changes at the leads would not be able to be detected by a reactance measurement, because the portion of the lead reactance on the total measured reactance is quite small.

Fig. 6: Mock-up transformer tested at KEMA Fig. 7: LV side with vertical LV lead exits In figure 6 the tank of the mock-up transformer at the short-circuit tests at KEMA site is shown and figure 7 shows a photo of the HV2 winding and LV lead exits of the mock-up taken at the visual in-tank inspection. The in-tank inspection of the core, HV winding, tank, bushings, lead exits and supporting structure gave no indication of apparent defects and changes in the mechanical condition that could impair the safe functioning of the transformer.

100⋅⋅ratedHVSC

HVSC

ÎÎ

8

5.2 Experimental destructive short circuit test on mock-up transformer (2nd test series) The purpose of a destructive short-circuit test is to identify the weakest point of the mechanical construction and to demonstrate the 20% safety margins for the most critical stress. The total number of tests was 14 along with additional set-up tests with reduced short- circuit currents. The test current was increased by a few percent after each test shot. The reactance change of the mock-up unit was measured before all tests and after each relevant test until a maximum reactance change of +1% could be recognized. This is also the upper limit allowed according to the IEC 60076-5. The tests were stopped at this level because it was most likely that mechanical movements in the mechanical structure had appeared. Figure 8 shows the reactance change of the test object after each current shot and the applied test current expressed in percentage of the reference current to obtain the 100% stresses. The first 3 current shots (5004-5007) with a maximum current level of 109.1% showed a stabilizing trend in the reactance change. The maximum impedance change was on a moderate level of +0.32 %. The reactance change showed a stabilized trend. The stresses in the mock-up transformer were (1.091² =1.19) +19% higher than the specified ones.

Fig. 8: Reactance changes and relative test currents at destructive tests Several SC-test experiments made in test laboratories showed that reactance changes of less than +0.5% are quite common and mostly not an indication of severe mechanical deformations. Such low level reactance changes are caused by a compacting effect on the windings and insulation structure. Impedance changes in the range of 0.5% to 1.0% can be an indication of progressive movements (see also CIGRE paper from KEMA [4]). The reactance change was higher than +0.5% after test number 5008 with a current level of 110.1%. The subsequent tests with increasing current level also led to a progressive increase of the reactance. The IEC limit of +1% reactance change was reached at a current level of 118.1%. The mock-up transformer was therefore stressed with forces which were 39 % higher than the specified and occurring stresses in the real 570 MVA GSU-transformer. Because of the progressive change of reactance between 0.5% and 1% reactance change it can be concluded that something happened in the transformer windings. Therefore it would be justified to derive a minimum safety margin of approximately 20% for the stresses from the tests. This can be concluded, when the test current of 110.1% after test number 5008 is considered as the critical current level. 5.2.1 Detected failure modes after dismantling of windings in factory After the short-circuit tests at KEMA, the mock-up unit was brought to the Siemens Weiz factory and un-tanked for visual inspections. The outside inspection of the active part gave no indication of deformations on the HV2 winding, the core, the clampings or the high-voltage and neutral lead structures (figure 9, 10). At the LVo vertical lead exit there was some buckling of CTC 1 (uppermost conductor) visible from the outside. The CTC

Relative variation of SC-Reactances during experimental SC-test

0,00

0,10

0,20

0,30

0,40

0,50

0,60

0,70

0,80

0,90

1,00

1,10

1,20

100426-5004

100426-5006

100426-5007

100426-5008

100426-5009

100426-5010

100426-5011

100426-5014

Test number

Mea

sure

d R

eact

ance

cha

nge

[%]

100

102

104

106

108

110

112

114

116

118

120Reactance change [%]Asym. SC current [%]

IEC-Limit

0,5% change

stabilizing trend

increasing trend in reactance change(progressive movement of LV conductors, likely)

[%]100⋅

⋅ratedHVSC

HVSC

ÎÎ

9

1 buckled near the wooden support probably caused by a thrust force from the LVo winding. All the other five CTC´s of the six in parallel showed no deformations (figure 11). After the upper common pressboard-ring was removed a tangential movement of the uppermost turn was observable as in figure 12. The uppermost LVo open turn tightens up to the insulation structure underneath and caused a tangential movement by pushing the vertical lead upwards.

Fig. 9: HV side with HV-leads Fig. 10: Flux return yoke and LV side with LV lead exits

Fig. 11a: LV lead exit( left view). Conductor buckling at LVo conductor 1 (right one)

Fig. 11b: Vertical LV lead exits (right view) Conductor buckling at LVo conductor 1

Fig. 12: LV winding top view (upper support ring dismantled). Tangential movement of LVo winding; LVi showed no movements

Fig. 13: LV winding (after HV2 winding dismantled). Tangential movement of upper LVo turns. Remarkable spiralling effect is visible by anti-clockwise upper turn rotation

10

This effect is more pronounced in the uppermost conductor (CTC1) because friction between the conductor and the spacers above and underneath plays an important role to prevent tangential movements. But friction in the conductor is basically a function of axial forces. The higher the axial force on the conductor turn, the higher the friction force to prevent tangential movements. Therefore, the so called “spiralling” effect is more pronounced at the uppermost part of the winding because friction is there lower compared to the winding centre. The cumulative effect of the axial compression force on spacers is highest at the winding centre. A tangential movement of the uppermost winding turn can be seen quite well in figure 13 by the non-vertical alignment of the axial winding sticks. A tangential movement was not observed at the lowermost turn because this turn was well fixed and supported at the lower mounting plate and layer transition supports. The innermost HV1 winding was also inspected and no abnormalities were detected. 5.2.2 Lessons learned from failure inspection In order to improve the withstand capability and prevent spiralling effects in helical and layer windings, the manufacturer has to take additional design measures. The most effective one would be to provide an infinite stiff radial winding support. With an infinite support, all radial forces would be transferred to the support underneath and there would be no tangential force in the uppermost and lowermost open turn. In practice this could be done by using optimized manufacturing methods and by reducing tolerances and radial shrinkage of insulation supports. Further, manufacturer shall be able to take into account friction at the winding turns by the evaluation of axial dynamic forces in windings. In the near future, additional efforts will be necessary in the transformer industry to model such effects with, for example, finite element methods and derive improved design and manufacturing measures to increase the overall withstand capability stress level of helical windings in GSU-transformers. 6. Sweep Frequency Response Analysis (SFRA) before and after short circuit tests

In figure 14 the sweep-frequency-response measurement results are given. The measurement was made in the “open circuit” method by applying the source between the HV and Neutral terminals, while the LV terminals were floating. The measured trace taken from the new mock-up before the first short-circuit test series is identical to the trace taken after the acceptance short-circuit tests. This finding correlates with the low reactance change of only +0.2% after the acceptance tests. The traces measured after the destructive short-circuit test series have a remarkable shift in magnitude and phase angle. Therefore it can be concluded that the windings did not experience mechanical deformations after the acceptance short- circuit tests.

Figure 14: SFRA plots taken at the mock-up transformer A second test with the “open circuit” method by applying the source between the LV terminals while the HV and Neutral terminals were floating also showed similar characteristics and sensitivity as mentioned above.

11

7. Conclusion EDF and Siemens took a completely new approach to demonstrate the short-circuit withstand capability of a large 570 MVA single-phase GSU- transformer. As a full-scale short-circuit test on the real unit was not feasible due to testing power limitations in high power laboratories, a mock-up transformer was designed, manufactured and finally short-circuit tested to demonstrate the mechanical performance. It was shown that it is possible to design a similar reference mock-up transformer according to well defined similarity criteria. The design of the windings was optimized in a way to demonstrate exactly the most important compressive hoop stress on the LV winding during the short-circuit tests. The benefit of such tests on a mock-up unit is that the manufacturer can also deduce the safety margins of the design by means of a destructive short-circuit test. This type of test is performed by gradually increasing current shots until the unit fails or experiences a notable increase in the winding reactance. A further advantage of such experimental tests is that the windings can be dismantled more easily compared to windings in real full-scaled units. Also final failure modes can be detected and corrective measures can be taken, when the tests have been finished before the manufacturing of the real unit starts. However, the design of a mock-up transformer can be very complex in order to achieve highest similarity to the real transformer unit. A lot of discussion is needed between the user and the manufacturer at an early stage (offer stage) because the cost of a mock-up may be very different depending on what is going to be tested and investigated. BIBLIOGRAPHY [1] IEC 60076-5 Edition 3.0 2006-2. “Power transformers-Part 5-Ability to withstand short circuit” [2] Electricité de France. “CST 72.C.042.02, Avril 2004 - Cahier des spécifications techniques” [3] Gerald Leber. “Short circuit test on a 440 MVA, 400kV GSU-Transformer” (Cigre Colloqium

Budapest 1999) [4] A.L.J. Janssen, L.H. te Paske.“ Short-circuit testing experience with large power transformers”

CIGRE Paper 12-105, Paris Session 2000

Documents

21, rue d’Artois, F-75008 PARIS A2-201 CIGRE 2014 http