Sheldrake, R

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A THYRATRON FOR THE NLC BASELINE MODULATORM. Parmenter, C.A. Pirrie, C.A. Roberts, P. W. Robinson and R. SheldrakeMarconi Applied Technologies Ltd., Waterhouse Lane, Chelmsford, CM1 2QU, UK.AbstractThere are several modulator techniques being investigated as candidates fordriving NLC klystrons in groups of between two and eight per modulator.The baseline design is a conventional line type modulator that would use athyratron as the switch to drive two klystrons. This paper describes athyratron designed to meet these operational requirements. The targetlifetime is 50,000 hours. Cost and reliability are of critical importance dueto the large number of modulators required. Several design decisions weretaken at the outset including cathode type, basic tube diameter and numberof high voltage gaps required. The development process includedmeasurement of cathode temperature as a function of average current andelectrostatic field analysis of the high voltage geometry. The rationaleunderpinning the design that resulted is explained. Since the specificationwas first defined, the pulse width has been doubled with a correspondingdoubling of the average current. The impact of this updated requirement onthe existing design is also discussed.1. INTRODUCTIONThe table shows the original acceptance test criteria, operational data and the upgradedrequirement as a result of the pulse width being doubled.  The most significant challenge is toachieve the target life of 50,000 hours; all other parameters have been demonstrated to be wellwithin the capabilities of modern thyratron technology.Parameter OriginalAcceptance Test Original OperationPresentOperationAnode voltage (kV) 90 75 75Peak anode current (kA) 10 7.5 7.5Inverse voltage (kV) 10 max not known not knownDi/dt (kA/us) 100 75 75Average anode current (A) 2.5 2 4Pulse duration (70%) (us) 2.2 2.2 4.4PRF (Hz) 180 180 180Jitter (ns) 1 1 1Run time or life (hours) 24 50,000 50,000Run time or life (shots) 1.55E+07 3.24E+10 3.24E+102. KEY DESIGN CONSIDERATIONS AND DESIGN DECISIONSThe work described here builds upon previous work [1,2] where the performance, life andsuitability of various thyratron types and designs were discussed. The more recent paper [2] discussesdata on the current life history of candidate tubes; it also identifies the optimum means of triggeringand considers critical circuit parameters, which need to be considered to maximise thyratron life. These are considered in more detail here.2.1 Cathode typeThe thermionic cathode is one of the key features that will determine the achievable lifetime.  Theprevious work indicates that the barium aluminate (BA) impregnated cathode size used for 4.5diameter metal envelope  tubes (e.g. CX1836 type) and the oxide cathode in the 6  diameter ceramictubes (e.g. CX2410 type) are the minimum cathode sizes that could achieve 50,000 hours life at 2Aaverage.Field data supports this conclusion for Marconi Applied Technologies BA cathode tubes [2,3]. Insufficient field lifetime data exists for the Marconi 6  oxide cathode tube at high (>2A) averagecurrent levels, but tubes of this type (CX2410) have been operational at KEK since 1995 at averagecurrents from 0.3 to 1.3A [4]. Recent data from KEK [5] shows that out of the 16 tubes received in1995 there have been three failures from poor connections to the cathode terminals and that theremaining operational tubes have accumulated an average service life of 17,000 hours to date (witha calculated MTBF of 85,000 hours!).  There are lifetime data for thyratrons of this type from othermanufacturers that indicate that several tens of thousands of hours are likely to be achieved undersimilar average current conditions.  Since it appears that there is little difference emerging in servicelife between the BA cathode in the 4.5  tube and the oxide cathode in the 6  tube, the oxide cathodeis preferred over the smaller BA counterpart simply on the grounds of cost.2.2 Reservoir typeA high capacity reservoir system has been selected so that gas pressure adjustments throughout tubelife are minimised.  The high capacity system is more complex than a conventional reservoir; thereare two reservoir capsules which operate at different temperatures and further auxiliary controlcomponents to compensate for input power fluctuations and ambient temperature variations. Thetotal gas content, however, is typically more than ten times greater than that of a typicalconventional reservoir.  This large capacity has the significant added benefit that the need forreservoir ranging  to compensate for gas clean-up losses is removed.2.3 High voltage structure and number of gapsPrevious work [1,2] has indicated that a minimum of three high voltage gaps are required for reliableoperation at 75kV under command or resonant charge conditions.  Two basic mechanical designs ofmulti-gap tubes exist as shown in Fig. 1. Fig. 1a is the stackable, repeatable, drift-space structure andFigs.1b and 1c are examples of two and three gap nested-cup structures, respectively.  On a partscount and therefore cost point of view, a design like Fig. 1b or 1c is preferred.2.4 High voltage structure materialsTraditionally, very high voltage structures have been faced with molybdenum because it has provedmore resistant to arcing than copper, especially at anode voltages above about 60kV. High voltagestructures above this level using copper have been built before but have tended to exhibit destructivearcing.  An inherent property of all multi-gap thyratron structures is the very high peak transientelectric fields that occur as a consequence of the way the tube commutates. The tube is triggered atthe cathode end and the high voltage gaps go into conduction starting with the gap nearest thecathode, followed by a delay of around 50 ns before the next gap, and so on.  The result is that thelast high voltage gap at the anode must withstand the full anode voltage for a period of around 50 nswithout arcing. Molybdenum appears to become electrically conditioned to withstand this (tubes havebeen operated successfully at up to 160 kV) but copper has been observed to suffer arc damage thatprevents this conditioning from occurring.  Despite this, copper would be preferred on the groundsof cost and since cost is such an important factor, the design of the high voltage structure wasundertaken on this basis.  This represented a risk worth taking.Fig. 2 is a field plot of a simplified representation of the top gap of the proposed prototypedesign and shows that there is peak electric field of around 13.6 kV/mm when a potential of 30 kVis applied.  With an applied anode voltage of 90 kV corresponding to the specification maximum,these figures are exceeded by a factor of three for a period of 50 nsec or so during commutation. Although the field during commutation is rather high at around 40 kV/mm, it is only a transient field. The limits for acceptable peak field strength under such transient conditions are not precisely known,but test results presented below indicate that they have not been exceeded in this instance.  The radiiof curvature on the selected geometry are larger than those previously utilised for copper high voltagegeometries as a result of the results obtained by undertaking the field plotting work.  The generallyaccepted maximum field strength for d.c. applied voltages in electron tubes with conditioned surfacesis around 10 kV/mm.  The field strength in the top gap with 25 kV applied (corresponding to the 75kV operating level) is around 11 kV/mm, but this field will only be present for a small fraction of asecond.  The significant magnitude of the transient applied peak fields can be reduced using arelatively simple circuit technique, as discussed next.                    (a) (b) (c)Fig. 1  Various thyratron high voltage structures.  Fig. 1(a) is a typical example of the stackable drift-spacemulti-gap structure, whilst figs. 1(b) and 1(c) are typical examples of two- and three- gap nested-cup geometries.Fig. 2. Electric field plot of the critical area within the proposed top gap geometry.  The z-axis represents theaxis of the thyratron .  Also shown are the trajectories of single electrons released from the surface of theelectrode below the anode.2.5 Saturating anode inductorTo reduce the anode heating arising from switch-on losses, and, more important in the light of theabove discussion, to reduce the voltage appearing across the top gap during thyratron commutation,it is proposed to use a saturating anode inductor with this tube design. Detailed work performed atCERN [6] indicates that a saturating anode inductor with a volt-second product of 1.6mVs will givea switch-on delay of approx. 70ns with a three gap of the drift space design and that this will preventany excess voltage build up across the top gap of such a tube. Shorter delays will offer proportionatelyless protection and it may be necessary to trade off the effect of the added saturated inductance onthe current rise time against the protection afforded to the top gap. Saturating anode inductors alsoafford some protection against the adverse effects of inverse voltages [7].3.0 FINAL PROTOTYPE DESIGNThe design that was selected for manufacture is shown in Fig. 3. The key features outlined above areincorporated: an oxide cathode from a 6  diameter thyratron, a high capacity reservoir system,copper electrodes and a nested-cup high voltage structure.  The tube envelope surrounding the cathodestructure uses copper in preference to the more conventional ceramic and the basic diameter of thecontrol grid and high voltage structures is 4.5  rather than 6 .  These features result in a hybrid designthat is tailored specifically for this application where cost, reliability and operating lifetime areparamount.AnodeRod (30 kV)ceramicElectrode below anode at 0kVPeak field on negativeelectrode is 13.6 kV/mmFig. 3  Drawing and photograph of final CX2437X design for 2A average current operation4. FACTORY MODULATOR TESTS4.1 Cathode Temperature MeasurementThe optimum cathode operating temperature is a balance between sufficient primary electronemission at one extreme and minimising thermal evaporation at the other.  Drawing from experienceand published data [8], this temperature lies somewhere in the range 750-780 OCB.  Thyratrons differfrom all other electron tubes in that cathode back heating by ion collection/impact will make asignificant contribution to overall cathode power dissipation.  A sample tube was manufactured witha sidearm window to enable cathode temperature to be measured with an optical pyrometer and fig.4 shows the measured cathode temperature against average current for the first CX2410 cathodedesign.  It can be seen that the quiescent non-operating cathode temperature is at the top end of theacceptable range and that the temperature climbs significantly as the average current switched isincreased (by increasing the prf).  Fig. 5 shows how cathode temperature varies with input heaterpower alone, from which it is deduced that the optimum input power is around 450W to achieve 750OCB. This equates to a heater current of 71A at 6.3V and the cathode heater incorporated into thedesign was selected to have these parameters.Given that a temperature difference of as little as +25 OC may double the rate of coatingevaporation [9] the importance of getting the temperature correct is evident.  Given this, it isperhaps surprising that the practice of operating thyratrons with reduced heater power to compensatefor back-heating has not been adopted much more widely.  The data in figs. 4 and 5 allow calculationof the required reduction in cathode input heater power to maintain the cathode at the optimumtemperature for operation at a given average current.Fig. 4 Cathode Temp. (OCB)  vs. AverageCurrent Fig. 5  Cathode Temp. (OCB) vs. Input Heater Power4.2 High Voltage OperationMarconi Applied Technologies does not have equipment to test an NLC tube under all conditionssimultaneously, but there are separate equipments that can test groups of parameters that collectivelyencompass most aspects of the specification.  The high voltage tests performed on two prototypetubes are summarised below. Both tests were carried out with a saturating inductor giving a switch-ondelay of about 18ns (Ferrite type CMD5005, 4 OD, 2 ID, 1  thick) [10].i) 85kV, 2.83kA, 0.6us pulse in a command charged (2ms with a 8ms hold), fast rate of riseof current (120kA/us) equipment, but at only 5Hz.ii) 70kV, 10kA, 6us pulse at up to 67Hz (4A average) in a resonant charged, slow rate of riseof current (8kA/us) modulator.Fig. 6 is an oscillogram of the anode current rise at 85kV with a reasonable gas pressure in the tube:10-90% risetime, including the overshoot, is 28ns, which gives a di/dt of 101kA/us. Jitter is 0.9ns700720740760780800820840350 375 400 425 450 475 500 525 550Input Heater PowerCathode Temperature7607808008208408608809000 0.5 1 1.5 2 2.5 3 3.5 4 4.5Average Anode CurrentCathode Temperaturepeak-to-peak.Fig. 6  Anode Current rise-time and JitterUnder high average current conditions in the 70kV modulator, the two prototype tubes, denoted withtype number CX2437, operated flawlessly.  One of the parameters measured was the rate of fall ofanode voltage, commonly called the dv/dt signal. The dv/dt signal provides a relative measurementof the gas pressure in the high voltage structure. Fig.7 shows the variation of the dv/dt signal as theaverage current is increased (byincreasing the prf) for the twoCX2437 tubes.  For purposes ofcomparison, similar data is shownfor the CX1536, which is astandard product used commonlyin high power linacs of varioustypes. It can be seen that theCX2437 thyratrons display alarger gas pressure variation thanthe CX1536 as average current isincreased. For operation at the2A average current condition ofthe original specification, theperformance of the CX2437design is deemed to be well withinacceptable limits. The differencebetween the performance of theCX2437 and the CX1536 is attributed to improved heat extraction from the surface of the anode dueto the larger diameter anode rod in the latter.   The average cross sectional area of the CX2437 anoderod is only 16% of that of the CX1536.With a pre-pulse of 10A on the grid 1 and a 0.5us delayed trigger pulse to the control grid 2,CX2437 jitter was around 0.9ns peak-to-peak under both sets operating conditions used in the factorytests and therefore within limits. If needs be, the jitter could be reduced further by using dc powersupplies to energise the cathode heaters or by increasing the rate of rise of voltage supplied by thecontrol grid 2 pulse generator.5. DEVELOPMENTS TO MEET THE LATEST SPECIFICATIONFig. 7   Dv/dt signal with average current00.511.522.533.540 2 4 6 8 10Average Currentdv/dtCX2437 s/n 5CX2437 s/n 6CX1536AFig.7 shows that the doubling of theaverage current to the 4A level causes asignificant fall in the internal gas pressurewithin the high voltage structure.  Thiscan be overcome simply by improvingthe heat extraction i.e. by incorporatinga larger diameter anode rod.  Fig. 8 showsa proposed modification where the anoderod has an average diameterapproximately four times that of the twoprototype samples. It is anticipated thatsuch a tube would exhibit performancesimilar to that of the CX1536 in fig.7 atup to 4A average current.   There is alsoscope within the existing envelope t oincorporate a cathode with a significantlylarger surface area, although the figuredoes not depict this.  A larger cathodewould be necessary to maintain the50,000 hours target life in light of thecathode back-heating increasingsignificantly at double the pulse width.ACKNOWLEDGEMENTSThe authors wish to thank the directorsof Marconi Applied Technologies forpermission to publish this paper and t othank those of our customers who havekindly provided life information for thispaper.REFERENCES[1] R. Sheldrake et al., EEVThyratrons for NLC KlystronModulators , Second Klystron-Modulator Workshop, SLAC, 1995. SLAC Report 481 March 1996.[2] C.A. Pirrie et al. Thyratron andModulator Design Considerations toMaximise Thyratron Life , 1998Klystron-Modulator Workshop, SLAC,1998.[3] Private communications on lifeinformation for CX1836A andCX2412A  types.Fig.8 Thyatron design with larger anode.[4] M. Akemoto et  al., Thyratron Performance in the KEK 8 GeV Linac , 25th LinearAccelerator Meeting.  SPring-8, July 2000.[5]Private Communication KEK, May 2000.[6] L. Ducimetiere & D. Fiander, Commutation Losses of a Multigap High Voltage        Thyratron , IEEE 19th Power Modulator Symposium 1990.[7] C.A. Pirrie et al.   Thyratron Design and Circuit Techniques to Overcome the AdverseEffects of High Inverse Voltages , 21st International Power Modulator Symposium, Los Angeles,1994.  (Also EEV Technical Reprint 195).[8] Research Study on Hydrogen Thyratrons , EG&G Inc, in three volumes as follows:Volume 1, S.T.Martin, S.Goldberg, 1953.Volume 2, S.Goldberg, 1956.Volume 3, S.Goldberg, D.F.Riley, 1957.[9] S.Goldberg, Cathode Phenomena and Life in Hydrogen Thyratrons , Proc. Seventh Symp. onThyratrons and Modulators, pp5-9, 1962.[10] Ceramic Magnetics Inc., published data.   http://www.cmi-ferrite.com

A Thyratron for the NLC Baseline Modulator.

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