Figure 1.0 TF Central Column Sliding Joint Fingers, Horizontal Legs Removed. THE TF FAULT

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1 Failure Analysis and Design Improvement of the Alcator C-Mod Toroidal Field Coil Sliding Joint Peter H. Titus, Stone & Webster Engineering Corporation, under contract to MIT Plasma Science and Fusion Center, 185 Albany Street, Cambridge MA Dr, Herb Becker, David Gwinn, Ken Rettman, Peter Stahle, MIT Plasma Science and Fusion Center Abstract-- In March of 1998 a short occurred in the TF system. Small amounts of copper and carbon tracking were found in the A-right TF. A crack in a finger in the central column was also found. Subsequent disassembly revealed arcing damage to the outer - upper horizontal leg finger joint region, at the lead connection leg where the voltage difference is large. Material was also removed from the neighboring torque shell. Significant degradation of the felt metal in the inner fingers at the central column was also found. This initiated re-analysis of the finger design, testing of the felt metal, and a decision to replace all the felt metal in the machine. Design improvements have been made but some uncertainty as to the cause of the failure remains. INTRODUCTION There are 20 TF picture frame coils in the C-Mod tokamak. Each coil has 6 turns for a total of 120 turns. All the turns are series connected and have a resistance of about 600 micro-ohms at LN2 temperature. Through the use of sliding joints, each of these picture frame coils is divided up into sub-components corresponding to each of the 4 sides. The TF magnet consists of 40 upper and lower horizontal arms, 20 vertical legs and the bonded central column assembly. Turn-to-turn insulation is G10CR and there are G10 housings into which LN2 cooling channels are cut. The TF coils weigh about 17700kg, or lbs. The toroidal field magnet assembly is described in more detail in [1] Alcator C-Mod uses sliding electrical joints in the corners of the picture frame TF coil. The intention is to isolate individual legs of the TF coil, and provide simple and direct mechanisms to support the primary loads in each of the legs. Inner leg centering loads are supported by wedging, or hoop compression. Horizontal legs transmit their vertical bursting loads to the large cover forg ings. The outboard leg transmits its radially outward loads to the outer shell, also called the torque shell. The concept simplifies the structure in a global sense, but substitutes delicate and complex sliding joints that require fairly complex analysis and careful fabrication. The sliding joints relieve the large thermal differentials resulting from the heat-up of the copper coils. The primary component of the joint is felt metal, made up of a mat of fine copper wires sintered to a copper backing. The compliance of the fibers improves contact. Each turn carries 250,000 amps at the 9 T TF design point. Each joint has four strips of felt metal, each about 18 cm^2 in area, yielding an average current density of 3.4 ka/cm^2. Electro-magnetic current diffusion can double the current density, and variations in contact area due to mechanical effects can increase it still further. Current density capacity of 9 kilo-amps per square centimeter has been demonstrated in recent (1998) small sample tests. The inner fingers of the central column are shown in figure 1. The assembly is represented in figure 2, in the form of a 2D finite element model used to study the joint's behavior. Figure 1.0 TF Central Column Sliding Joint Fingers, Horizontal Legs Removed. THE TF FAULT In March of 1998 a short occurred in the TF system. There was an arc across the main TF first to last turn, which connects to the bus, -the point of maximum voltage difference. The arc blew a ½ golf ball size gouge in the outer cylinder, with similar sized damage in the TF coil. There is a recess cut across the face of the outer TF leg that was meant for LN2 distribution. Some of the recess close to the finger had exposed copper and allowed collection of material that could cause an electrical bridge. This was the area where the arc occurred. Later

2 inspections yielded a possible source of the material that may have caused the arc. When the cover came off there were indications of copper sputtering on the G-10 shroud above the inner fingers. Copper dust was evident. Under microscopic examination, a portion of the dust appeared to be fine melted copper beads. A crack was observed on the top of one of the TF central column fingers. After complete disassembly, the cracked corner came off the finger. When the coils were disassembled there was substantial damage to the upper-inner joints. There was degradation of the felt metal, notches eaten out of the fingers of the horizontal leg about 1/3 up from the plasma side, and copper deposited on the mating fingers of the central column. After further disassembly of the machine. The lower finger joints were not found to be damaged. There was some small indication of etching of the lower central column fingers, but minimal degradation of the felt metal. This initiated re-analysis of the finger design, testing of the felt metal, and a decision to replace all the felt metal in the machine. TF FAULT WHAT COULD HAVE CAUSED IT? a. An error in the initial design The analysis was re-visited. Variations in the physical parameters that characterize the model were investigated. Worst case configurations were investigated. The design was found to be sound b.operating conditions different than the original design Actual operating PF scenarios, disruption scenarios, different TF charging profiles, all were checked. No problems were found. c.material flaws in the felt metal Flaws were possible but difficult to determine from FM from the machine, and samples from original manufacture were not available. d.spring plate assemblies yielded or were damaged? No plastic deformations of the spring plate chevron heights were found. Measured force deflection was nominal. e.were the springs inserted with the correct interference? Were the slot measurements accurate? Measurements were made with G-10 gauges that did not expand the fingers. Measurements of the force needed to remove the springs showed some springs were below nominal pressure. There was no correlation between worst damaged felt metal and poor spring plate pressure. But there is still some uncertainty in this evaluation. UP-DOWN ASYMMETRY WHAT CAUSED IT? Damage occurred on the upper fingers. Current polarity is a possibility. The cathodic - anodic relationship of the FM contact is different on top and bottom. However, reversal of current in felt metal tests showed no change. But no field was applied in these tests. Cryostat pressure is maintained above atmospheric. It was unlikely but possible that low nitrogen pressure in the cryostat and air intrusion could have caused damage to the upper felt metal pads. Assembly errors are a possibility. Different crews assembled the top and bottom, but a review of the procedures and experience with the assembly produced no obvious differences. Cooling between shots favors the lower portion of the machine, but analyses of the temperature difference, discussed later in this paper did not yield excessive temperatures. ACTIONS TAKEN TO DETERMINE THE CAUSE Metallurgical evaluation of the Felt Metal and dust was performed. Force-deflection measurements of the felt metal were made. Magnetic, structural, and thermal reanalysis of the fingers was performed. Bench testing the felt metal, cyclic testing of the felt metal in an automated test fixture was conducted. Bench testing of the spring plates, and force-deflection measurements of the spring plates in the PSFC material lab was performed. A review of MAST experience and design was done. ACTIONS TAKEN TO IMPROVE PERFORMANCE New felt metal was installed all around the machine. Dimensional surveys formed the basis for spring insertion, a better gauge was developed to check the slot opening. This had an expanding slot gauge using opposed ramps or saw teeth with a hand grip actuator and a dial guage. It could apply a small pressure on the slot opening to remove the "slop" before the slot size was measured. A new spring plate design was developed for the inner finger joints. It produced a higher pressure, more compliance, and allowed plastic deformation to match the slot openings. Better specs on the felt metal, its silver plating, and its sintering. were used. Use of an Electrodag coating on the felt metal was bench tested, and used at assembly. It has been shown to improve the "settling-in" of the felt metal Felt metal for the repair/reconstruction of the finger joints in September and October of 1998 was better documented in terms of the process controls used during manufacture than in the initial assembly of the machine. The felt metal is in the fully annealed condition when sent to the plater. The silver plating thickness was increased. Plating was performed in accordance with ASTM procedures Vol Des. B322-85, B and B700-81, and a documented procedure was established. This

3 consisted of a reverse current at 50 amps in Potassium Hydroxide, a spray rinse, three successive rinses in DI water, a bath in an activator for 30 seconds(20% sulfuric acid), again three successive rinses in DI water, and ultrasonic cleaning, a silver strike of 20 amps for 1 minute, a silver plate at 7 amps for 55 minutes followed by more rinses, ultrasonic cleaning in water and methanol and finally a high pressure HEPA filter air dry. Improved insulation was used in the notch area where the arc occurred. Liquid Nitrogen feeds to this area was supplied with its own filter. Thermocouples were added near the fingers and in the central column. MECHANICAL ANALYSES, SUMMARY OF RESULTS A single TF coil blade and finger set is modeled in this analysis. A 2-D model, shown in figure 2, is used and represents a slice through the vertical and horizontal legs of the TF coil just above the corner radius. Both in-plane and out-of plane Lorenz loads are applied. These are calculated by "embedding" the 2D model in a Biot-Savart model of the poloidal coil system. Early design analyses used.13 N/mm^3 applied as an inertia load. This was derived from a MAP current diffusion analysis based on 9T and 3MA Ip operation. For the recent Biot Savart calculations, local currents in the finger were taken from Pillisbury's MAP analyses. These were Figure 4. Lorentz Forces on 2D Finger Model, 30 degree Angle, Fourth Current Set, skn4.inp essentially confirmed by Myatt's analyses [4]. Gap elements are used to model the spring plates and felt metal. Separate elements model the stainless steel and copper layers of the TF coil. Displacements of the model boundaries are taken from the analysis of the TF inner leg/column. A simple 3D solid model of a plate was used to show that the three dimensional torsional behavior of the finger coils could be safely neglected. Currents were taken from shot # , which was a high OH1 current shot (20 ka), with an 8 T toroidal field.and a.841 MA plasma. Upward and downward disruption shots # , and # were considered as well. The latest analysis runs are summarized In the table below. A nominal spring interference of 0.01 inches produces a non-uniform pressure on the felt metal of between 100 to 800 psi. the average is about 400 psi. Table 1 Analysis Runs Run # Sprg Int (in.) Sprg Int (in.) Felt Metal Material Modeling Elastic-up to Elastic-up to Elastic-up to Elastic-up to 1500 Comments Nominal spring plate interference (.01 inner.01 outer) Based on Production Samples Table 2 Nominal Behavior - With Multiple loading Run # 734, Felt Metal Pressures (MPa) Inner Spring Interference =.01, Outer Spring Interference =.01, Nominal Lorentz Loads (in Parenthesis is Approx % Area of Contact Lost) Loading Inner Felt Metal Min Inner Felt Metal Outer Felt Metal Min Outer Felt Metal Max Max Init Spring Lorentz st Lorentz mm (15%) (10%) After 1 st Load (20%) 2 nd Lorentz mm (15%) (3%) After 2 nd Load (20%) Behavior of Second Loading Not Much Different than the First Errors or variations in current diffusion analyses could produce significant changes in out of plane (OOP) loading of the fingers. The table below shows the possibility of significant loading variations with only 15 changes in the orientation of the currents with respect to the poloidal field vectors. Table 3 Effect of Inclination of the Current Vectors Lorentz Load - 9 Tesla, bunching factor = 2.0 PF Fields from Shot# Inclination of Total Load in Newton/mm^3

4 2D Model Finger Region (N) Loading Table 4 Summary of Contact Loss for Load Cases Considered Loaded Loss of Contact Unloaded Loss of Contact Nominal 15% 20% Inner/Outer Variation 20% 15% Too Low Spring Pressure 20% 25% Larger OOP Displacement 30% 25% Plasma Reversal 20% 25% Twice the Lorentz Loads 30% 20% Poor Sizing of Spring Packs error in slot measurement, or "slop" not measurable with gauge. 60% 75% There could be some combination of off-normal loads that would worsen in-service behavior, but installation problems, particularly with slot measurements, look most suspect. Thermal analysis discussed later shows temperatures start to rise significantly with 50% or more loss in contact. THERMAL/ELECTRICAL BEHAVIOR OF THE FINGER JOINT Micro Ohmcm^2.8 sec sec sec sec sec sec sec Variations in resistivity alone do not produce excessive temperatures. There was some concern that cooldown between pulses might not be complete, and reach the 90 K limit imposed on the subsequent shot. In figure 8, a 9 Tesla shot with a 120 K start temperature, and the worst FM resistivity would reach unacceptable temperatures. In the table below, a median resistivity would require a start temperature of 160 K before there was a problem. Table 6 Effect of Start Temperature on Felt Metal Temperature K Full Contact of FM Pads, K=5.59W/m/degK, 15 Micro Ohm-cm^2 Start Temp,.8 sec 2.5 sec 4.8 sec 6.0 sec deg K Two analyses are available to estimate the temperature of the felt metal during a pulse. The first is a one-d transient conduction model, and the second is an ANSYS 2D model. The original thermal calculations were done by Table 7 Felt Metal Temperature Figure 6, 2D Transient Conduction Model Besen and Gwinn using the MAP code. The recent input is the same as previously used, based on Pillsbury s current diffusion model, ref[2]. Felt metal properties are from the MAST report [3]. There is some possibility of variation in the thermal properties of the FM. Tables 5 through 11 and Figures 7,8, and 9 represent transient conduction simulation results for a variety of variations from the nominal conditions. Table 5 Effect of Felt Metal Resistivity on Felt Metal Temperature, K Full Contact of FM Pads, 9 T, K=5.59W/m/degK, 80 deg K Start Temp FM Area Resistance /3 (5mm out of 15mm)Contact of Outer FM Pad, K=5.59W/m/degK, 80 deg K Start Temp FM Area

5 Resistance µ Ω-cm^2.8 sec sec sec sec sec Table 11Felt Metal Temperature (Inner or Outer Pad) - Effect of Thermal Conductivity, Full Contact of Inner and Outer FM Pad, 10 Micro Ohm - cm^2 80 deg. K Start Temp FM K W/mdegK sec sec sec sec sec For thermal problems to develop, significant loss of contact is needed even for a wide variety of thermal and electrical resistivities. This, along with the possible loss in contact due to assembly procedures leads to the conclusion that improved methods of spring insertion are needed. IMPROVED SPRING DESIGN Table 8 Felt Metal Temperature Inner vs. Outer Loss of Contact K=5.59W/m/degK, 80 K Start Temp, 15 Micro Ohm cm^2 Inner 2/3 Loss Outer 2/3 Loss.8 sec sec sec sec sec sec sec Table 9 Felt Metal Temperature Inner vs. Outer Loss of Contact K=5.59W/m/degK, 80 K Start Temp, 10 Micro Ohm cm^2 Inner 2/3 Loss Outer 2/3 Loss.8 sec sec sec sec sec sec sec Table 10 Felt Metal Temperature (Inner or Outer Pad) - Effect of Thermal Conductivity 1/3 (5mm out of 15mm)Contact of Outer FM Pad, 10 Micro Ohm-cm^2, 80 deg K Start Temp. Typically Max Temp is in Outer Reduced Pad Early in the Pulse and in the Inner Pad Later in the Pulse FM K W/mdegK.8 sec sec sec sec sec There are two goals for and improved spring pack design. The first is the capability to operate at a higher pressure than 400 psi. This reduces the sensitivity of the joint to increases in Lorentz forces. The second is to maintain this pressure over wider variations in the distances between the fingers, whether derived from tolerance build up or felt metal compaction. Figure 10 4 Plate Spring Assembly Figure 11 4 Plate Spring Calculated Pressure vs. Deflection Originally in the central column joints there were 2 spring plates, each.035 in thick and.5 in wide. A compression of the pack of.02 inches yields a pressure of about 750 psi. If you could stack two of these pairs of springs you could double the height change, halving the sensitivity to assembly tolerances and/or felt metal shrinkage. If they were fully flattened the stack height would be.14 in If you reduce the thickness of the carrier, and work the springs closer to their fully flattened condition, there is almost room for 4 springs instead of 2.

6 In practice the thickness of the spring plates had to be reduced to.032 in thick to allow insertion in the slots, which average.145 inches wide. In Fig. 11 that yields a pressure of about 800 psi. Pressures increase very quickly for smaller slot openings. Insertion methods were worked out in a slot mock-up, shown in figure 12. The 0.08 inch extension shown in figure 10, centers the pack radially in the slot. been analyzed. The original analysis used plastic behavior beginning at 800 psi. Measured force deflection for the October 1998 production material showed yielding behavior at about 1500 psi. Test data is plotted in figuire 14. The 2D analytic model was re-run, with the stiffer FM and higher spring plate pressure, from the new 4 plate springs, modeled as well. Felt metal remained in contact, as evidenced by the toroidal displacement compatibility shown in Fig 13, along with resulting felt metal pressures. Cyclic tests of the Felt Metal to 9kA and 1500 psi in an automated INSTRON test fixture are in process. Figure 12 Four Plate Finger Spring being Inserted into the Slot Mock-Up. Insertion of the new spring packs was simulated in a mock-up to ensure successful assembly. A survey of the as-fabricated felt metal repair was done and slot openings computed and compared with the mock-up slot opening. With consideration of compliance of the felt metal, all the slots were expected to allow insertion of the new spring packs. This was confirmed at assembly. Fitting the fingers together in C-Mod is difficult. Alignment and slot tolerances of.001 inch are needed. The horizontal leg needs to be handled with felt metal in place. Damaging felt metal during assembly with the central column is a possibility. ANALYSIS WITH PRODUCTION FM PROPERTIES Finger response for the final measured force deflection behavior of the felt metal installed in the machine has Figure 14 Felt Metal Pressure vs. Deflection CONCLUSION The sliding joint failure in March 1998 is still not fully understood. For a nominal assembly, there is some possibility that worst case combinations of the uncertainties in current diffusion, spring plate preload, felt metal compliance and other parameters might be marginal for the 3MA 9 Tesla design point, but not for the 1 MA 8 Tesla peak operation that C-Mod had experienced. Loose fit of the spring packs showed the largest effect on contact loss. Simulation of the thermal behavior of the joint yielded no excessive temperatures, for expected variations in resistance and conduction properties. While loss in contact had to be significant before temperatures became unacceptable, imprecise spring pack insertion could produce this degree of contact loss. Mechanical design of the springs was improved, and slot measurement and spring pack assembly procedures were improved. REFERENCES [1] Alcator C-Mod Toroidal Field Magnet Assembly" W. Beck 14 th SOFE, Oct , San Diego, page 292 [2] Memo to D.B. Montgomery, From R.D.Pillsbury Jan CMOD-TF Transient Thermal Analysis of the C-Mod TF Coil Sliding Joint [3] Measurement of Feltmetal Material Properties MAST (Meg Amp Sherical Tokamak) Design Note Feb G.Voss, AEA Technology Space and Defense Systems Department, Culham England, [4] 3D Coupled Electromagnetic, Thermal Current Diffusion in the Finger Joints of the Alcator C-Mod Toroidal Field Coils R.L.Myatt, P.H.Titus, 17 th IEEE SOFE, October 1997, San Diego California

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