R. Viswanathan Electric Power Research Institute,

Palo Alto, Calif. 94303

^

V**

Materials for Generator Retaining Rings Retaining rings are among the highest stressed components of electricial generators, whose integrity is critical to the reliable and safe operation of the generators. The steel currently used for this application is susceptible to stress corrosion cracking and is also limited by yield strength to about 1200 MPa. The use of alternate materials with improved mechanical properties can enhance the reliability, ef­ficiency, size capability and availability of generators. This paper is a state-of-the-art review of the materials technology pertaining to retaining rings.

Introduction

Retaining rings are massive steel rings that are utilized to retain the circumferential conductor coils wound around the two ends of the generator rotor. These rings are subject to large centrifugal forces which increase with generator size. They are currently made of an 18Mn-5Cr steel which is cold expanded, an operation which both forms the ring and work-hardens it to the desired strength level. There are three problems with this current practice. It does not now appear feasible to develop a yield strength much greater than 1200 MPa in an 18Mn-5Cr retaining ring. This strength limitation places an upper limit on the size of generators. While the 18Mn-5Cr steel has good fracture toughness in air, it is susceptible to stress corrosion cracking in water, damp hydrogen, chlorides and other aggressive environments. This susceptibility is believed to have been a major cause of several retaining ring failures. The need for cold-expanding the alloy to strength requires a complex and delicate manufacturing procedure. Manufacturing difficulties have, therefore, resulted in limited sources of supply, difficulties in procurement, and extended forced outages. The objective of this paper is to present an overall perspective of this industry wide problem by a review and consolidation of past developments in the field, leading into potential alternatives for the future.

Retaining Ring Failures

A major fraction of all the structural failures of generator rotors have been caused by the failure of the retaining rings. The failure of a retaining ring may be defined as "the inability of the ring to perform its intended function due to cracking or due to complete fracture, thereby warranting its replacement." Sources of damage have been identified as overheating in installation or removal for inspection, mechanical damage due to improper lifting, electrical burn­ing, and cracking [1]. Should damage from any of the above sources be serious enough to render the retaining rings un­satisfactory for further service, a utility would be faced with

Contributed by the Materials Division for publication in the JOURNAL OF ENGINEERING MATERIALS AND TECHNOLOGY. Manuscript received by the

Materials Division, October 15, 1980.

an extended forced outage, if replacement retaining rings could not be readily obtained. At present, retaining rings are long lead items and procurement delays of six months to one year are not unusual [1, 2]. The economic consequence of the complete fracture of a retaining ring can be catastrophic. In the case of the Skaerback failure, for instance, pieces from the fractured retaining ring were thrown out through the stator end winding and stator housing into the turbine room [3]. Single pieces penetrated the thick concrete wall into the control room and destroyed two control panels for another generator. The total loss for the power company alone was estimated to be $1,800,000, not to mention the costs of repair of the extensive damages to the generator which was still within the warranty period [3].

A list of retaining ring failures has been published recently by Hagaman [4]. Based on this and on additional information obtained by private communications, the summary in­formation provided in Table 1 has been prepared. The number of failures reported in the table is expected to be conservative since replacement of cracked retaining rings is done sometimes in the field without widespread reporting. Retaining ring failures are also known to have occurred in India, New Zealand and other countries but have not received worldwide attention. The data in Table 1 show failures regardless of generator ratings, speeds, materials used and the manufacturer.

The most publicized failures of retaining rings occurred in two 100 MW units in Toronto in April 1954 [5, 6]. In both rings fracture had started at mid-thickness in the ventilation hole nearest the tapered end and had spread in each direction. The cracks were intergranular and there was no evidence of environmental contribution to the cracking or of fatigue. Hardness tests made on the internal surfaces of the ven-tillation holes showed hardness values in the range 300-657

Table 1 Summary Information Relating to Retaining Ring Failures Number of known ring failures 30 Number of catastrophic failures leading to generator damage 6 Number of ring failures in 3000-3600 rpm machines 20 Number of ring failures in machines rated > 500 Mw 8 Number of ring failures during 1970-1980 14 Number of ring failures related to U. S. manufacturers 13 Number of failures in H2 O cooled generators 2

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DPH compared to only 280-340 DPH at regions away from the hole indicating severe strain hardening of the inside surfaces of the holes. The maximum hardening was observed near the mid thickness, where the cracks started. It was concluded that the failures were caused by a combination of high stresses, stress concentration near the holes and strain hardening and associated low ductility inside the hole surface. An identical failure with similar contributing causes had occurred earlier in an English electric unit resulting in destruction of the entire rotor, but the information had not been widely disseminated until after Toronto failures [7], As a result of these experiences generator manufacturer modified their ring designs to eliminate ventilation holes.

Presence of local stress concentration and strain hardening due to machining were also cited to be contributing causes to the catastrophic failure of a NiCrMoV ferritic steel ring in 1974 at the Nanticoke station of Ontario Hydro [8]. The ring had failed by a single fast transgranular fracture in the radial axial plane, propagating axially from the shrink fit at­tachment end. The fracture origin was found to be in the radius of the bayonet lug used for fitting the ring. Microhardness measurements showed the surface layers to be at least 50 percent harder than the body of the rings. Since the fracture toughness of the ring in general was excellent, it was concluded that failure occurred by a combination of localized plastic strain due to stress concentration, surface hardening resulting from machining and embrittlement due to the cooling hydrogen.

The earliest recorded instance of nonmagnetic steel retaining ring failures due to stress corrosion cracking was observed in 1951 in England [5]. The problem was ex­perienced in a 25 MW generator which had been built in 1944 and had been crated and stored. When the machine was uncrated in 1951, two large axial cracks were found in one of the retaining rings. The cracks were attributed to stress corrosion due to a combination of exposure to salt air and the presence of tensile shrink fit stresses. Some finished rings which had been crated and stored for shipment were also found to have developed transgranular cracks presumably due to residual machining stresses acting in concert with a chloride environment. Numerous stress corrosion failures attributed to nitrates, cooling water and other environments have been reported since [9-13]. Two major ring failures in water cooled generators have been reported recently [4, 12, 13]. At Skaerback, in a 250 MW (330 MVA) machine of completely water cooled design, the retaining ring on the exciter side burst into four parts. Shortly before the damage occurred, the only sign was an increase in the vibration level. The cause of the damage stemmed from the fact that, during a long period out of operation, the remains of cooling water led to stress corrosion cracking, at an incompletely protected location of the ring in the region of the shrink-fit sealing at the end ring. At Barseback, the failure is also suspected to be due to stress corrosion in presence of cooling water.

In summary, failures of nonmagnetic retaining rings in service have been attributed to one of two major causes; one of them is the brittle fracture at regions of stress con­centration due to a combination of high local stresses, and the presence of work hardened, low ductility material. The other major cause of nonmagnetic ring failures is stress corrosion cracking induced by chlorides, nitrates, moisture or cooling water. The major cause of magnetic ring failures is cited to be brittle fracture at regions of high stress concentration aided by hydrogen environment. Presence of temper embrittlement in the steel was also found to have contributed to the fracture in some instances. Table 2 summarizes some of the key features of retaining ring failures which have been reported in some detail in the literature.

Most of the retaining ring failures reported to date have occurred by intergranular cracking. Some exceptions to this,

however, have been noted. The cracking mode in the Parsons' retaining ring that cracked in storage due to chloride con­tamination (Item 1, Table 2) was transgranular [5]. Hydrogen assisted cracking of the Ontario hydro-Nanticoke ring oc­curred by transgranular cleavage [8]. Speidel found that stress corrosion cracking of 18 Mn-5Cr steel can be intergranular as well as transgranular in distilled water [14]. Stress corrosion cracks have been found to follow deformation bands as well as grain boundaries in the evaluations conducted by Fritz and DeForest [15]. It is generally known that chloride induced stress corrosion cracks in austenitic steels can be trans­granular or intergranular. It is also known that in ferritic steels hydrogen assisted cracking can be trans­granular or intergranular depending on the yield strength and the degree of temper embrittlement present. In view of all this conflicting evidence it is often impossible to determine the failure mechanism based on the fracture path.

Material Property Requirements

Table 3 provides a list of the principal material properties that have to be considered for the retaining ring application. They are discussed in detail below:

(i) Physical Properties The use of magnetic rings on a rotor results in a greater magnetic end flux leakage with resulting extra heating in the stator coil ends and iron losses in the end region of the core. Additional excitation is required to compensate for this leakage and the total machine efficiency is reduced. The temperature rise occurring in a magnetizable retaining ring during operation may also endanger the shrink fit of the ring and cause damage to the insulation. Magnetic rings made of ferritic steels were commonly employed in the past since they offered many other advantages. They were made of inexpensive low alloy steels. Since the steels are heat treatable, any desired strength level could be achieved, without the need for the complex cold work processes in­volved in the manufacture of nonmagnetic rings. Fabrication of the magnetic rings was easier and the mechanical properties tended to be more uniform through the wall and more isotropic. Despite these advantages, nonmagnetic austenitic type rings have gained widespread acceptance over the years since the efficiency losses due to the use of magnetic rings have become unacceptable with the increased power ratings of generators. In addition to a low magnetic permeability, the retaining ring materials should have high thermal and elec­trical conductivities to prevent overheating and to maintain high efficiency. Since assembly of the rotor is performed by shrink fitting of the ring onto the rotor body, the ring material should have a reasonably high coefficient of thermal expansion. This minimizes the temperature required for the shrink fit and avoids damage to the ring and to the electrical insulation. A low density is desirable, since for a given size of the ring, the centrifugal stresses due to the ring are lower. A list of physical properties typical of the 18Mn-5Cr alloy in use today is contained in Table 4 [16].

(ii) Mechanical Properties. The diameter of retaining rings is determined mainly by the generator rotor diameter. Over the past few decades the generator capacity has been in­creasing steadily, warranting a corresponding increase in the rotor diameter and hence the diameter of retaining rings. This means an increase in the hoop stress in the ring since it is estimated that for a 3600 rpm machine, at 20 percent over-speed 68 percent of the ring stress is caused by the weight of the steel ring itself. Required stengths of retaining rings have been such that the 20 percent overspeed average hoop stress is equal to about 70 percent of the 0.2 percent offset yield strength [17]. Large 2-pole generators in the U.S. reached a plateau of about 1000 MW (1120 MVA) some years ago and the next step to 1300-1400 MW may follow. One of the major

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Table 2 Details of major retaining ring failures

Year of Year of Rating/ Failure Commission Utility Manufacturer Speed

1951 Ontario Hydro. R. L. Hearn Unit #1

1952 Ontario Hydro. R. L. Learn Unit #2

British English Electric Electric Authority Fulham Station

1946 Vaal South Afr

1971 Denmark Brown Skaerback-3 Boveri

Hydro Nanticoke 2

1978 1976 West Germany KWU Scholven H

1979 1975 Sweden ASEA Barsebaek 1

20 MW/ 1500 rpm

250 MW/ 115 3000 rpni

500 MW/ 3600 rpm

8Mn-8Ni-4Cr

8Mn-BNi-4Cr

Yield r e n g t h , F a i l u r e Diagnosii

100 MW/ 150 1800 rpm

60 MW/ 100 8Mn-8Ni-4Cr 3000 rpm

714 MW/ 3000 rpm

8Mn-8Ni-4Cr

HiCrMoV magnetic

Rings cracked in stor­age. Transgranular, ace cracking due to chloride.

Catastrophic brittle frac­ture. Intergranularj cracking due to stress concentration at ventil­ation holes, work hard­ened metal at the bore of holes and residual stresse:

Catastrophic brittle frac­ture in test tunnel during rotor balancing at normal Speed. Intergranular frac­ture similar to Ontario Hydn Failures.

Catastrophic brittle frac­ture. Intergranular fractun due to SCC by Nitrates.

Catastrophic brittle frac­ture in water cooled gen­erator. Intergranular fracture due to SCC by water

Catast

high :

>phic brittle frac-Intergranular frac-

ssisted by hydrogen, ted at a region of

Ring fractured without extensive damage to gen­erator. Intergranular SCC cracking due to moisture exposure during shipping, extended storage, or cold standby period.

Catastrophic fracture -similar to Skaerbaek failure. Water cooled generator.

UD—Undisclosed

Table 3 Material considerations in retaining rings

A.

B .

C .

D .

Magnetic

Conductive

i ) E l e c t r i c a l

i i ) Thermal

Densi ty

Thermal Expansion

A.

S .

C .

D.

E .

F .

E l a s t i c Modulus

Yield S t r eng th

T e n s i l e S t reng th

F rac tu re Toughness

in a i r (K I C)

Fa t igue crack growth

r a t e (da/dn)

S t r e s s Relaxa t ion

A.

B .

F r a c t u r e toughness (K I S C C )

a n d r a t e of crack growth

( d a / d t )

i )

i i )

i i i )

i v )

Hydrogen

Mois ture

S a l t water

Other environments

encountered in

manufac ture , s t o r a g e ,

sh ipp ing or o p e r a t i o n

Temperature in t h e

ran. ge - 40 to +55°C

A.

B .

C .

Formab i l i t y

I n s p e c t a b i l i t y

A b i l i t y t o

a t t a c h to

the r o t o r

Table 4 Physical properties of 18Mn-5Cr steel

Density

Specific heat

Thermal Conductivity

Coefficient of thermal expansion at 100°C 400°C

Magnetic Permeability

Specific Electrical Resistivity

Young's Modulus

7 .

0 .

30

14 19

1 .

7 6

2 .

8 5

5 8

. 8

. 0

. 0

g/cm"

J / g ' >C

W/m °C

x 10 ' X 10

0 0 5

yf t -cm

0 3 X 10 !

-v "V

•C

•C

3 MPa

limits on increased unit size is imposed by the retaining ring where stresses already approach the yield strength of 18Mn-5Cr alloy. In turn, this restricts the rotor slot depth and rotor diameter, forcing field currents to increase to achieve ampere turn loadings and providing new problems in rotor thermal capability. At the present time, however, there is no great

interest on the part of U.S. utilities to build large units and the emphasis is on reliability, availability, and efficiency of the generator. It must also be emphasized that the availability of higher strength ring materials is not the only issue impacting the manufacture of larger diameter rotors. Many other issues may control the decision - such as deflections, copper

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Table 5 Stress intensity (K|) data for typical flaws

Type

Type

Type

Type

Type

of

1

2

3

4

Flaw

- Short, deep surface crack

c = 2a

- Long, shallow surface crack

c > 10a

- Internal spherical crack

c = a

- Internal, elliptical crack

c = 2a

Initial Flaw

size a, mm

.25

.51

.25

.51

.25

.51

.25

.51

KI, HPa /m

22

31

27

39

14

21

20

28

operating stresses, rotor forging materials availability and trade offs, wedge stresses, assembly technology, rotor length, etc. In addition to enabling construction of generators of higher ratings, any improvement in retaining ring yield strength may allow for efficiency gains in the generator for all ratings of 2-pole machines due to the ability to accommodate increased copper loadings in the rotor. This again is a very complex issue and has to be considered in the context of many other factors such as magnetic capabilities, ventilation requirements, electrical creepage, copper operating stresses, exciter limits and other factors.

Retaining rings are currently available with guaranteed 0.2 percent yield strength level at mid wall of 1200 MPa. A yield strength requirement of 1400 MPa for the immediate future and of 1600 MPa for the next generation of retaining ring materials is not unrealistic, as one of the design options.

In addition to a high yield strength, it is desirable that the retaining ring materials have a high elastic modulus to resist deflection under load. A major challenge to the achievement of higher yield strength in retaining ring materials is that of maintaining reasonable fracture toughness in air and in adverse environments. Prior to the recent advances in fracture mechanics technology, the value of absorbed energy at room temperature based on Charpy tests was used as a qualitative index of the toughness of the steels. In the last few years more quantitative approaches based on linear elastic fracture mechanics have become available and are used as a basis of design.

A simplified fracture mechanics calculation can be per­formed, using the stress intensity expressions given by Tiffany and Masters [18], to illustrate the methodology for estimating the fracture toughness requirements. It is assumed that the flaw has the worst possible orientation, i.e., the major plane of the flaw is normal to the hoop stress. The hoop stress is assumed to be 840 MPa, corresponding to 70 percent of the yield strength. Two different flaw depths of 0.25 mm (0.01 in.) and 0.5 mm (0.02 in.) are considered. The stress intensity (Kj) values corresponding to four different types of possible flaws have been calculated and are listed in Table 5.

From Table 5, it is evident that the Type 2 flaw (long surface crack, c > 10a) represents the worst case that has to be considered for fracture mechanics analysis. It is also reasonable to assume that surface flaws of this type of about 0.5 mm depth represent the limit of detectability by current ultrasonic techniques.' Hence a Kj value of about 40 MPa Vm is the approximate lower limits of KIC (critical stress intensity for unstable crack propagation) that is acceptable for the steel for the conditions specified. If local regions of stress con­centrations such as holes and keyways exist, then the requirements on the K[C will be higher. Magnetic ring materials used in the past had very high initial values of K l c , with the exception that embrittlement phenomena such as temper embrittlement could degrade the KIC to dangerously low values. Commercial nonmagnetic ring materials in use

The limit of detectability can vary widely depending upon the capability of the laboratory, metallurgical structure of the ring, proximity to geometric discontinuities and many other factors.

±110

100

i : 30

20

10 I i I mill i i mini I I MINI I I I I I I I I I I I I I I I I I I I I I mil

103 104 105 106 107 10B 109

Number of Cycles

Fig. 1 Stress versus cyclic life diagrams for 18Mn-5Cr steels. The data enclosed by the scatter bands is for 145 ksi yield strength steel [11]. The triangular data points are for a 130 ksi yield strength steel [13].

today have K lc values well in excess of 100 MPa Vm even at high yield strength levels, see Table 6. Hence, in the absence of subcritical crack growth mechanisms such as fatigue, stress corrosion or hydrogen embrittlement, fracture toughness in air per se is not expected to be a limiting factor, provided that local regions of stress concentration are avoided by careful design.

Fatigue strength under completely reversed bending stress is another important material property. A peak load generator that is started and stopped twice per day during its 40-year expected life is subjected to about 30,000 load cycles. Based on published S-N curves, for 18Mn-5Cr material of about 1015 MPa yield strength the tolerable fatigue stress range for notched bar specimens for a 30,000 cycle life is about 500 MPa in air, and for smooth bar specimens is about 700 MPa [16], Based on the ratio of 106 cycle fatigue strength to the ultimate tensile strength, obtained from the data in Fig. 1 it can be estimated that a 1400 MPa yield strength level, a fatigue strength of about 550 MPa may be adequate. In the past, fatigue fractures have not posed any serious problems. Presence of a corrosive environment, however, may alter this picture substantially. Fatigue crack growth type data are generally unavailable. Detailed stress analysis information, for a specific design would also be needed before the crack growth data can be utilized. For illustrative purposes, assuming that the alternating stress is as high as the maximum hoop stress of 840 MPa (for a 1200 MPa yield strength steel) and utilizing a crack growth rate of 15 micro inches/cycle [19], it can be shown that a type 2 preexisting flaw of depth 0.5 mm (0.02 in.) will not grow to critical size in less than about 40,000 cycles. In actual practice, the fatigue stresses are expected to be considerably below the endurance limit of the steel and fatigue crack growth rates would be markedly lower. Failure of retaining rings by fatigue in air has, therefore, never been a major concern.

Since most generator designs utilize a shrink fit to keep the retaining ring securely attached to the rotor, the retaining ring materials must have a high resistance to stress relaxation by creep at the operating temperatures. Grain size control is essential for inspectability, since the detectability of flaws by ultrasonic techniques can be affected by interfering effects due to grain size. The ASTM grain size 0 for instance, causes an attenuation of sound of approximately 0.15 dB/mm. Normal production routes for nonmagnetic rings produce material with grain sizes in the ASTM range of 0-4, with a desirable range of 2 to 3.

Another metallurgical requirement is that the ring should not be susceptible to degradation phenomena, such as temper embrittlement or sensitization, during the processing or assembly of the ring. In magnetic rings of the Ni-Cr-Mo-V type, slow cooling, or tempering in the range 375-575°C can lead to temper embrittlement. Acting in concert with hydrogen embrittlement temper embrittlement, can reduce the

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Table 6 Summary of fracture toughness on retaining ring alloys

Yield Type Basic Alloy strength Environmental Composition MPa Test

1BMn-5Cr . 5 C - . 1 N i - . 5 S i

18Hn-5Cr . 5 C - . 7 S 1 - . 0 9 V

Aged A286

Fe-16Mn-18Cr Cold worked

50%

19Mn-6Cr .5C- .2V . 5 S i - . 2 N i

18Mn-6.5Cr . 6 C - . 7 S 1 . 14V- .2Ni

J9Mn-8.2Cr 3Ho-.9V (WRETAINS a l l o y , Table 4-1)

19Mn-8.2Cr 3Mo-0.9V (WRETAINS a l l o y , Table 4-1)

34Ni 3Ti 3Ta e t c ( a l l o y T, Table 4-1)

34Ni-2 .5Ti -7 .5Ta e t c ( a l l o y M, Table 4-1)

1200

1125-

1125 1200

1380

1240

1225

t365

1325

1400

1370

MPa fm

182 (plane -(stress)

other Type of Information

da 5x10 cm/ d t s e c a t 90°C

F a i l e d i n 4 h o u r s

—

^ 1 = 1 0 6 c m / r , e c d t a t 90»C

F a i l e d i n 3-14 d a y s

u n a f f e c t e d

14

2 2

2 2

165+

r a d i a l 141 +

c i r c u m -e r e n t i a l

1 3 7

r a d i a l 112+ c i r c u m ­f e r e n t i a l

r a d i a l 43 c i r c u m f e r ­e n t i a l 23+

For r a d i a l o r i e n t a t i o n 1000 hr f a i l u r e a t K = 31

_-.

For r a d i a l o r i e n t a t i o n 1000 h r f a i l u r e a t K = 7 2

r a d i a l 9 9

113 da/dN identical in H-, and

•Threshold value of K below which crack propagation was not observed, in presence of an environment other than dry air.

Approximately 12 x 6 x 1cm fatigue precracked WOL specimens taken from rings, wedge loaded to a predetermined K.. Cracking in nitrates intergranular. In distilled water mixed mode. 2.5 x 2.5 x .3cm specimen stressed to 630 MPa as a beam. Exposed to 20% Nacl salt spray at 40-C in a cabinet or . between 5% Nacl and 100% humidity till a crack could be seen. Single edge notched specimens, precracked by fatigue, then subjected to stepwise rising load.

loaded as cantilevers and time to failure determined ring. Precracked compact tension specimens, subjects

Precracked specimei Samples from a 122< 350 kPa H2S environments. Samples from a 64cm ring. Procedures same as in 5. Samples from laboratory heat, cold worked and aged. Procedui Samples from laboratory heatf sustained load up to 560 MPa ii

d to function of Initial K. rising load, distilled water, 560 kPa H,

Procedure for H-S test described in text.

fracture toughness ( t̂hreshold)of these steels to values as low as 20 MPa Vm [20]. In the case of nonmagnetic austenitic rings, sensitization of the grain boundaries due to carbide precipitation can occur during slow cooling from the solution temperature through the critical range of about 538-760°C [21] and lead to reduced fracture toughness, particularly in coarse-grained material. Special attention must, therefore, be paid for obtaining the best possible quench. Inadvertent overheating into the critical temperature range during assembly is known to have caused premature stress corrosion cracking in some nonmagnetic retaining rings in the field.

(Hi) Environmental Considerations. A majority of the retaining ring failures to date have been attributed to stress corrosion type of cracking. The stress required can arise from the residual stresses from cold fabrication of the rings, machining stresses, shrink fit assembly stresses or operating stresses. The aggressive environments causing cracking may

include lubricating oils and fluids used during machining, solvents and paints used for cleaning, dye penetrant in­spection, condensed moisture, salt water and industrial air environments encountered during storage and shipping and hydrogen or water used for generator cooling. Residual stresses from fabrication are alone sometimes sufficient to cause cracking of the rings as evidenced by ring cracking in storage. Since retaining rings are at present manufactured only in France, Germany and Japan, transportation of the rings over long distances, especially by ocean freight is inevitable. Environmental considerations relating to storage have assumed major importance particularly in light of today's economic and regulatory climate that has resulted in long term storage of generation equipment.

Published data relating to the resistance of retaining ring type alloys to environmental cracking are meager. A summary of the available information is included in Table 6. The test procedure used, specimen geometries and evaluation criteria

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K (ksi N in)

20 40 60 80 100 120

10-'°

i r 23 C (73 F)

90 C (194 F)

_L_^ L 0 20 40 60 80 100 120 140

K (MPa x m)

Fig. 2 Stress corrosion crack growth data for 18Mn-5Cr steel in aerated distilled water at 23 and 90 C [16]

vary from one investigation to another. The type of data deemed most desirable for evaluating the risk of en­vironmental failure are the threshold stress intensity for crack propagation (KISCC) and the rate of crack growth (da/dt) as a function of the applied stress intensity. Each of these parameters by itself is inadequate, as may be illustrated from the data of Speidel, see Fig. 2 [14]. Comparing the K1SCC results alone at 23°C and 90°C, one might conclude that the susceptibility of 18Mn-5Cr steel to cracking in distilled water is the same under both conditions. The rate of crack growth however, is higher at 23°C at K > 80 MPa vm. Similarly, addition of halides and nitrates to distilled water were shown to increase the crack growth rate substantially, without altering the KISCC value. These results show that in any fracture mechanics analysis both the K1SCC and the da/dt data should be considered for the specific operational conditions (temperature, stress level and environment).

The effect of specimen orientation on the time to failure for a 1250 MPa yield strength retaining ring material is illustrated in Fig. 3 [19]. Specimens were notched and precracked, loaded as cantilevers and the time to failure was determined as a function of the initial applied stress intensity K. The test procedure is a nonstandard procedure and is currently nder study by the round robin test program on stress corrosion by ASTM E24.01 and G1.06 task group. The results of Fig. 3 clearly show that the LC orientation in which the notch was in the circumferential direction resulted in shorter lives com­pared to the CL orientation in which the notch was in the radial-axial plane. For both orientations the synthetic sea water environment was more deleterious. Threshold K values as low as 10 MPa Vm are obtained, indicating the high susceptibility of 18Mn-5Cr to environmentally induced cracking. The preference for crack propagation in the cir­cumferential direction is also confirmed by the results of Hull [17] and by observations on service failures of rings. For instance, in the Scholven failure, the principal initiating stress corrosion crack was circumferential, which then ruptured under overload in the axial direction [11].

Considerable judgment has to be exercised in applying the data from laboratory tests to the service performance of materials. For instance, rising load tests, in which the K1SCC values are measured under a monotonically increasing load may yield KISCC numbers that are nonconservative due to the fact that the test does not allow for incubation phenomena. The rising load test can therefore be used for very high-

o • Synthetic sea water

® m Distilled water

LC CL

J I I I I I I I I L I I I I I I J L_L 10 100 1000 6000

Time to Failure (hr)

Fig. 3 Effect of environment and specimen orientation on the time-to-failure of 18Mn-SCr steel [19]

strength materials or for medium strength materials tested under severe environmental conditions. For evaluating materials of reasonable ductility under moderately severe conditions, constant deflection tests or other slower test methods have to be employed even though they may be cumbersome and time-consuming. Superimposition of cyclic loading even at low frequencies may modify the static en­vironmental crack growth behavior substantially. In con­ducting tests under simulated environmental conditions, careful control of the environment must be exercised. For instance, tests to evaluate resistance to cracking in hydrogen must use hydrogen gas of a purity that is the same as that used as the generator coolant, since trace amounts of impurities like oxygen or moisture may alter the crack growth behavior. The severe anisotropy and location dependence of the crack growth susceptibility in actual retaining rings must also be taken into account in conducting the laboratory tests.

The general methodology used in evaluating the influence of the environment on the risk of retaining ring failure can be illustrated as follows using the data of Speidel [14]. Since the Kiscc =7 MPa Vm in distilled water, even the smallest detectable flaws will grow subcritically at a hoop stress of 840 MPa.2 Using the stress intensity expression applicable to a type 2 surface crack for a reported KIC value of 146 MPa Vm the critical crack size aa for brittle fracture can be calculated to be approximately 0.7 cm. Using the reported value of crack rate growth of 10~6 cm/s at 23°C, corresponding to an initial K level of 39 MPa (type 2 flaw 0.5 mm, see Table 5) the time for onset of unstable brittle fracture can be calculated to be about 200 hr. Presence of chlorides would accelerate the crack growth rate further and lead to failure in fewer hours.

In comparison with distilled water, high humidity air and chlorides, hydrogen environments appear to be generally less aggressive. The limited amount of data available indicate that the KISCC values in hydrogen are much higher than those obtained in aqueous environments. There have been no reported failures of 18Mn-5Cr rings attributed to hydrogen in hydrogen cooled generators. Crack growth data in hydrogen are not available to do detailed fracture mechanics type calculations.

Admittedly, the fracture mechanics calculations used for illustration purposes are simplistic and involve assumptions regarding the flaw shape, acuity, and distribution as well as the operating conditions. The calculations also treat only the growth of single straight cracks, rather than the case of several cracks growing simultaneously or that of branched cracks. In spite of these limitations, the calculations show the tremendous influence of adverse environments on the fracture

70 percent yield strength for a 1200 MPa yield strength steel.

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Table 7 Typical compositions of nonmagnetic retaining ring alloys

Alloy Desiqnat ion

Commercial SNi-SMn

-4Cr

1BMn-4Cr

18Mn-5Cr-V

18Mn-5Cr -V-N

IBMn-lBCr

Developmental Modified

A286

W-RETAINS

EPRI Berkeley Alloy T

EPRI Berkeley Alloy M

Approximate year of

I n t roduc t i on

1920

192B

1954

1967

1978

1968

0.5 0.B

0.65 0.75

0.06 0.53

C r Mn

Method of Ring

Manufactu:

4.0 8.5 7.5 Bal Hot-cold forging during early years; then cold

4.0 - 18.0 Bal Cold Expansion

0.14 0.1 6.2 1.5 19.0

2.2 0.01 0.005 1.25 0.3 0.26 14.5 25.5 1.2 Bal Cold expans.

3.03 0.01 0.07 0.95 8.2 0.59 19.8 Bal

0.01 1.0 0.5

0.01 1,0

0.3 5.0 34.5

,3 5.0 34.0

lal Expected to be hot forgec

240

200

— 180

S "» T3

2 140

OJ

o 120

100

80

8Ni-8Mn-4Cr compositions

-

/ \

EPRI-Berkeley alloys T, M /

W-retains ® alloy 9

18Mn-5Cr v modifications , . compositions . . y4

—>p >p * p - N modifications

I I I I I

—

:

— _

-

—

1500

1400

1300

1200

1100

1000

900

800

700

600

Fig. 4 Yield strength requirement of retaining ring steels as a function of year of manufacture

behavior of currently used retaining ring materials. In fact, susceptibility of retaining rings to environmental cracking may well be the single most important limitation in the utilization of water cooling for generator rotors and in the future development of higher rated hydrogen cooled generators.

Evolution of Ring Materials and Recent Developments

The general trends in the composition and yield strength requirements of nonmagnetic retaining ring alloys are illustrated by the data in Table 7 and in Fig. 4. Retaining rings made until about 1938 were based on the composition 8Ni-8Mn-4Cr-0.5C. During the decade of the 1920's, the rings were made by a hot-cold forging operation. Due to difficulties in keeping the deformation uniform, the rings exhibited widely varying mechanical properties over the circumference. The problem was overcome by use of a cold expansion technique. The first cold expanded rings were produced in the late 1920's. Due to demand for material with yield strengths approximating 900 MPa and due to wartime shortages of nickel during the late 1930's, complete substitution of nickel by manganese was affected and rings with the basic com­position of 18Mn-5Cr-0.5C3 were made. Efforts to increase the yield strength further led to experimentation with ad-

The chronology of alloys discussed here is based on the experience of Krupp MetaJ-Und Schmiedewerke as described in reference [16]. There is some ad­ditional information indicating that cold expanded rings of the 18Mn-5Cr composition were manufactured as early as 1928 at Poldi in Choslovakia.

ditions of vanadium and columbium. Current compositions containing vandaium and nitrogen have enabled achievement of guaranteed yield strength levels as high as 1200 MPa. At this yield strength the alloys based on the 18Mn-5Cr com­position seem to have reached the limit of their capability.

Significant developments of alternate compositions by generator manufacturers during the 1970's have consisted mainly of the development of A286 type alloys by General Electric [22] and the development of the WRETAINS alloy by the Westinghouse Electric Corporation [17]. The A286 rings made clearly demonstrated the potential for achieving yield strength levels up to about 1200 MPa coupled with good stress corrosion resistance. Large scale use of the material, however, was not pursued due to disadvantages in terms of economy, reproducibility, availability and producibility of rings relative to the standard retaining ring alloy. The WRETAINS alloy has demonstrated potential for yield strength levels ap­proaching 1365 MPa with fracture toughness and stress corrosion properties that appear to be better than those of the 18Mn-5Cr alloy. Commercial utilization of the alloy may become possible if rings of the alloy can be manufactured successfully on a consistent basis.

To improve resistance to environmentally assisted cracking, ring manufacturers have also developed alternate 18Mn compositions with chromium levels increases to 18 percent. Preliminary evaluations of rings made by one of the ring manufacturers using an 18Mn-18Cr-0.5N alloy have shown that the alloy has excellent fracture toughness in air and resistance to cracking in a variety of environments [25]. The alloy has an inherently higher strength prior to cold work so that a given strength level can be achieved with less cold work compared to the 18 Mn-5Cr alloy [26]. Consequently, the cost of the rings made from the new alloy is expected to be only slightly higher, despite the higher chromium content. The degree of anisotropy in mechanical properties is also expected to be lower. Some concerns, however, remain which may have to be resolved before the 18Mn-18Cr gains widespread ac­ceptance. The yield strength of the alloy very closely ap­proaches the ultimate strength, posing a limitation in some generator designs. Both tensile strength and yield strength have been observed to decrease with temperature in the range 150-600°C. The room temperature strength has also been found to decrease due to prior exposure in the above tem­perature range [27]. Evidence of severe localized pitting has been found in concentrated chloride environments [27]. Susceptibility to grain boundary sensitization due to exposure in the range 550-800°C has been observed [28]. Pending further characterization of the 18Mn-18Cr steel. Some

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Table 8 Physical properties of EPRI-Berkeley alloy (alloy T) Density (lb/in3) 0.16 0.20

Density:

Thermal Expansion coeff ic ien t :

Res i s t i v i t y :

Magnetic Suscep t ib i l i t y :

E las t i c Modulus:

8.26 gm/cm

15.8 x 106 /'a

72.7 pfi-cm

< 1.01

1.9 x 105 MPa

generator manufacturers have cautiously moved toward gaining in-service experience on a limited scale by utilizing rings made from the steel in their smaller machines. Current versions of the 18Mn-18Cr steel are also limited to a yield strength of about 1050 MPa. Research is in progress to achieve higher yield strength by increasing the nitrogen content of the steel.

There are many disadvantages associated with the current fabrication practice for Mn-Cr steels. In view of the costly and sophisticated tooling and experience required for cold expansion of the rings there is only a limited number of ring suppliers and extended delays in procurement are sometimes encountered. Cold expansion also results in variation of the mechanical properties across the thickness of the ring and the variation becomes very pronounced at high yield strength. For instance for a 105 cm diameter 7.5 cm thick wall ring cold expanded to a mid wall yield strength of 1240 MPa the ID-OD yield strength may range from 1450 to 1100 MPa [15]. In addition, the mechanical properties are anisotropic. The rapid strain hardening capability associated with the cold ex­pandable alloys can prove to be detrimental if machining is carried out without extreme care. Localized strain hardened regions of low ductility, including the formation of strain-induced martensite can occur, under these conditions and can promote brittle fracture.

To overcome the problems associated with current alloys and fabrication practice, EPRI initiated a program of alloy development research at the University of California at Berkeley. The objective of this research was the development of a high-strength nonmagnetic austenitic steel which would be resistant to environmentally assisted cracking under the imposed service conditions and which would achieve a yield strength of 1400 MPa or greater in the heat treated condition while retaining reasonable fracture toughness. This objective has now been statisfied and an alloy, which has for con­venience been designated alloy T, has been developed in laboratory-size heats [24]. Alloy T is an iron base alloy (see Table 7). Since it contains no carbon, it is not susceptible to sensitization. The strength is derived primarily from y' (Ni3Al,Ti,Ta) precipitation. After forging and suitable heat treatment, the alloy develops a yield strength of 1435 MPa while retaining fracture toughness in air of 110 MPa Vm. The alloy has an ultimate tensile strength of 1575 MPa, elongation of 16 percent and reduction of area of 53 percent in room temperature tensile tests. The Charpy absorbed energy at room temperature is about 27 Joules. The tensile ductility and impact toughness level compare very favorably with those of 18Mn-5Cr steel evaluated at much lower yield strength level.

A list of the important physical properties of alloy T is furnished in Table 8. Comparison of this data with the data for 18Mn-5Cr (Table 4) shows that the properties of alloy T are comparable to that of the current alloy.

Notch bar tensile tests have shown that the alloy has no tendency for premature failure in hydrogen. The fatigue crack growth rate of the alloy in hydrogen (1 atm) is identical to that in air at room temperature. Sustained load tests using precracked compact tension specimens stressed to a K level of

1

°\

1 1 x _ 7 7 i n

S Glass

^ 64 in Graphite/Epoxy

\ 60 in \ Boron/Aluminum

33 in Betillium

B

1

•

n 4 3 i n u Al Alloy 7075

1 1

1 1 1 1

52 in ^ B e / T i

53 in ^T i -6AI -4V

46 in 1 18Mn-5Cr Steel o I

Density (g/cm3)

Source: Weslinghouse Eleclric Corporation

Fig. 5 Allowable stresses and ring sizes as a function of density for candidate retaining ring materials

88 MPa Vm show no tendency for subcritical crack growth in a 3 percent NaCH solution for test times up to 1 month.

Precracked compact tension specimens of the alloy have been subjected to subcritical crack growth testing in two environments: (1) 560 kPa H2 and (2) 350 kPa H2S. The test procedure was more severe than the conventional rising load test, in that load and displacement were applied cyclically in a series of increasing steps until clear evidence of crack growth was obtained; The remainder of the test consisted of holding a constant displacement during continued growth of the crack under conditions of decreasing stress intensity K. The values of stress intensity at crack arrest were estimated to be about 80 MPa Vm in both H2 and H2S.

In the course of the EPRI investigation, a second class of alloys have also been developed, collectively designated as alloy M, which can achieve the target yield strength after annealing and aging. These alloys are modifications of alloy T and utilize a higher content of hardening element to induce precipitation of a tetragonal y" in addition to 7 ' . Alloys of the M class are, however, not immediately suitable for manufacture of retaining rings, since the fracture toughness exhibited by these alloys at the 1400 MPa yield strength level appears inadequate. Further research is under way to optimize the composition and heat treatment parameters of alloys T and M to improve their strength-toughness characteristics. A concurrent program for large scale engineering evaluation of alloy T is also under way.

No discussion on retaining ring materials can be complete without some consideration of alternate materials with at­tractive strength-to-density ratios. Since a major portion of the stress on the ring is due to the weight of the ring itself, higher strength-to-density ratios will enable larger attainable stresses and hence permit use of larger diameter rotors. Figure 5 illustrates the relation between density and allowable stress for several potential candidate materials. Among the many materials mentioned in the figure only the titanium alloys and the Graphite/Epoxy composite material have received serious attention.

Rings made from titanium alloys are believed to be in commercial operation in some generators in the USSR. The technological barriers to the widespread commercial use of Ti-6A1-4V type alloys appear to be the lack of data regarding degradation of the alloys in generator type hydrogen en­vironments and the possibility of low temperature stress relaxation. The use of titanium alloys in shrink fitted assemblies may also require design modifications because the elastic modulus and the coefficient of thermal expansion are different from those of currently used materials. A project to

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evaluate titanium alloys for retaining rings is under con­sideration at the EPRI.

Composite materials offer many advantages over metals for the retaining ring application. Unlike metals, they have no yield point and behave elastically up to fracture. They do not creep under stress. They are generally immune to hydrogen attack and are nonmagnetic. They possess high stiffness and strength-to-density ratios. They are widely available. Composite materials of the graphite/Epoxy type have been in wide use for many structural applications in the aerospace industry. The material and design data available have therefore been derived primarily from experience on section thicknesses up to 1.25 cm. Since composite retaining rings will require thicknesses of up to 8.75 cm, material design data must be developed which will reliably predict the performance of such thick structures in air and other pertinent en­vironments. Moist environments can cause degradation of the structure. In view of the very low coefficient thermal ex­pansion, assembly of the ring by shrink fitting on to the rotor is not feasible. Alternate means of fastening the ring have to be explored. The low electrical and thermal conductivity of graphite composites may also require more design modifications in terms of minimizing thermal heating due to negative sequence currents. Any inadvertent overheating would readily damage the rings. Most of these areas are currently being investigated by EPRI. With the present state of development, utilization of materials other than steel remains a long range prospect. For the immediate future, further development and commercialization of the newer steel compositions that have been developed in recent years appear to hold considerable promise.

Acknowledgments

I am grateful to Messrs. A. F. Armor, J. Edmonds, R. I. Jaffee, and D. N. Poole of EPRI for reviewing the manuscript and for providing useful input to the report. Discussions with Mr. R. T. Hagaman of the Westinghouse Electric Corp., D. R. DeForest of General Electric Co. and J. Joyce of Utility Power Corp. are also gratefully acknowledged.

References

1 Stanley, R. J., Detroit Edison Co., 2000 Second Ave., Detroit, Mich. 48226, 1980, Personal Communication.

2 "On-the-Ball Operating and M & R Teams Avert Major Crisis," Report of Retaining Rings Failure at Penn Electric's Warren Station, Electrical World, Aug. 1,1978.

3 "Faulty Retaining Ring Possible Cause to the Failure at the Skaerback Station," Danish Engineers Weekly, May 15, 1973.

4 Hagaman, R. T., "Failure Experience with Generator Rotors," Proceedings of the EPRI Workshop on Rotor Forgings for Turbines and Generators, Palo Alto, Sept. 14-17, 1980.

5 "Failure of Forged End Bells on Large Electric Generators," Metal Progress, ASM, July 1956, Vol. 70, No. 1, 1954, p. 69.

6 Gibb, C , "Report on Investigation into the Failure of two 100 MW Turbo Generators," Proc. of the Inst, of Mechanical Engineers, Vol. 169, No. 29, 1954, p. 511.

7 Vickers, V. J., Discussion of item 5 on p. 528 of the same report. 8 Jolly, C. B., Murphy, M. C , Paterson, A. N., and Petty, D. J., "The

Failure of an End Ring of a Generator Rotor Operating in a Hydrogen Atomsphere," presented at CEA Thermal and Nuclear Power Section, Halifax, Nova Scotia, Oct. 1975.

9 Lissner, O., "Stress Corrosion in Austenitic Steels by Nitrates," The Engineers Digest, Dec. 1957, Vol. 18, No. 12, p. 571.

10 Lissner. O., "An Interesting Case of Stress Corrosion," ASEA Journal, 1957, Vol. 30, No. 5, p. 85.

11 "Report of the Failure of a Rotor Retaining Ring of the Scholven 'H ' Generator," Kraftwork Union report, M-TE 76434/1, 1978.

12 Kugler, H,, "Damage to Turbogenerators," Der Maschinenschaden, Vol. 49, 1976, No. 6, pp. 221/235. English translation available from Allianz Technische Information, D-8000, Munich 44, P.O. Box 24.

13 "Reports Tell How 610 MW ASEA Generator Exploded, How 335 MW Unit Shorted," Electrical Week, May 28, 1979, p. 7.

14 Speidel, M. O., "Stress Corrosion Cracking in Fe-Mn-Cr Alloys," Corrosion, 1976, Vol. 32, No. 5, p. 187.

15 Fritz, K. C , and DeForest, D. R., "Progress in the Development of Higher Strength Non Magnetic Retaining Rings," Journal of Materials, JMLSA, Vol. 4, No. 3, 1969, p. 647.

16 "Retaining Rings," Product Literature, Krupp, Metal-Und Schmie-dewerke, Altendorfer Str. 104, D4300 Essen, Germany.

17 Hull, F. C , "High Strength Austenitic Nonmagnetic Alloy," U.S. Patent 4121953, Oct. 24, 1978.

18 Tiffany, C. F., and Masters, N. J., "Applied Fracture Mechanics," ASTMSTP381, 1964, p. 249.

19 Ohhashi, J., and Sawada, S., Japan Steel Works, Ltd., Muroran, Japan, 1980, Personal Communication.

20 Viswanathan, R., and Hudak, S. J., "The Effect of Impurities and Strength Level on Hydrogen Induced Cracking in 4340 Steels in Effect of Hydrogen on Behavior of Materials," A. W. Thompson and I. M. Bernstein, Eds., p. 262, 1975, TheMetaiiurgicai Society of AIME.

21 Kohl, H., Wekstoffe undKorossion, Vol. 14, 1963, p. 831. 22 Fritz, K. E., and DeForest, D. R., "High Strength Turbo Generator

Retaining Ring Forgings of an Age Hardenable Austenitic Alloy, Journal of Metals, JMLSA, Vol. 3, No. 3, Sept. 1968, p. 629.

23 Benson, R. B., Jr., Kim, D. K., Atteridge, D., and Gerberich, W. W., "The Resistance to Embrittlement by a Hydrogen Environment of Selected High Strength Iron Manganese Alloys," in conference on stress corrosion cracking and Hydrogen embrittlement, Unieux-Firminy, France, June 10-16, 1973.

24 Chang, K. M., and Morris, J. W., "High Strength Austenitic Alloys for Generator Retaining Rings," EPRI report CS1808, Apr., 81.

25 Bellows, R. H., and Gibbs, E., presented at the "P900" colloquium, Oct. 23, 1980, Krupp Stahl, Essen, Germany.

26 Heringdorf, H., Stein, H., and Heinrich, H., Ibid. 27 McKintyre, P., Ibid. 28 Unneberg, L., Ibid.

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