SURFACE HARDENED NAVAL MARINE GEARS WITH REFERENCE TO ALTERNATIVE MEANS OF SURFACE HARDENING

ROGER BARKER & GEORGE C. MUDD

SURFACE HARDENED NA VAL MAMNE GEARS WITH REFERENCE TO

ALTERNATIVE MEANS OF SURFACE HARDENING

THE AUTHORS

Roger Barker joined David Brown Gear Industries Limited in 1965 as a Development Metallurgist after graduating from Manchester University with a B.Sc. (Hons) degree in metallurgv. Progressing to Works Metallurgist, then Assistant Chief Metallurgist, he was appointed Chief Metallurgist in 1978.

George C . Mudd started his career in 1951 in design, moving in 1953 to research and development with particular emphasis on the fundamental nature of gear tooth performance in all ap- plications. In 1964 he became Chief Gear Test Engineering Manager at David Brown Gear Industries, Huddersfield, England, and has held management positions within the engineering function becoming Director of Engineering in 1978. He holds B.Sc., C. Eng., and MI. mechanical engineering degrees.

ABSTRACT

It has become common practice in naval marine gear units of European manufacture to take advantage of the greater load carrying capacity resulting from a surface hardening process. The surface hardening processes open to the gear designer are many and vaned, and each has advantages and disadvantages.

This paper examines the three principal applicable processes, explains the characteristics of each and how the disadvantages may be controlled. The load carrying capacity of gears made with the different surface treatments is then discussed, including the effects of hardness gradient, residual stress and size on capacity.

INTRODUCTION

T H E DEPLOYMENT OF SURFACE HARDENING TECHNIQUES IN naval marine gear units has been common practice now for a good many years in Europe. The reasons are simple - greatly improved fatigue resistance means increased gear load carrying capacity which in turn leads to significant reductions in size and weight, two vitally important considerations for naval marine gear units.

Figure 1 provides a graphic example of the order of size difference when comparing a through hardened gear, UTS 60170 t.s.i. with a case hardened gear, both required to transmit the same power.

In this paper we examine the three principal surface hardening processes applicable to naval marine gear units, namely: carburise case hardening, nitriding, and induction hardening. The paper is in two parts, the first of which identifies the important metallurgical quality and process control characteristics of each process and outlines measures which must be adopted to ensure that the required quality is achieved. The second part then

252 Naval Engineers Journal, May 1984

explains the way in which the properties developed are related to allowable load carrying capacity for the particular process and gear tooth geometry.

HARDENING PROCESSES

CARBURISE CASE HARDENING

General Characteristics

In the context of marine gears this commonly involves three separate operations: carburise and slow cool, re- heat and oil quench, and temper. Carburising, usually in the temperature range 900-950°C, in the presence of a high carbon-potential gaseous atmosphere, involves diffusion of carbon into the steel surface. The depth of carbon penetration and therefore case depth is highly controllable over an extremely wide range.

Re-heat and oil quench transforms the high carbon case to hard martensite, typically of the order 61-63Rc, whilst the bulk core material transforms to some much lower strength value dependant on base steel composition and ruling section. (Figure 2) Due to the different transformation characteristics of the case and core, the quenching process induces a residual stress profile with favourable compressive stresses at and near the surface within the case, changing to balancing tensile stresses in the region of the case and core interface and within the core. (Figure 3) The tempering process changes the brittle as-hardened, martensite to a tougher form with some loss in hardness to a minimum of 58Rc and some modification of the residual stress profile, reducing both the compressive and tensile peaks.

This relatively simple profile of the process masks a multitude of detailed features which must be considered and controlled during carburise case hardening. These features affect the actual load carrying capacity achieved or the practical and commercial feasibility of the process and are discussed in the following sections.

Atmosphere Control

In the first place, tight control of gas carburising atmosphere composition is vital to prevent:

a) Excessive austenite retention (Figure 4) b) Excessive carbide precipitation (Figure 5 ) c) Low carbon leading to soft non-martensitic transfor-

d) Localised surface decarburisation (Figure 7) mation products (Figure 6)

BARKER/MUDD SURFACE HARDENING OF MARINE GEARS

Figure 1. Case Hardened V's Direct Hardened Gear Same Transmitted Power.

D e p t h Below Surface ( ~ n s ) 1 Ggure 2. Typical Hardness Gradient-Carburise Case

Hardening.

I

Figure 3. Typical Residual Stress Distribution-Carburise Case Hardening.

I Figure 4. Excessive Retained Austenite.

All these features can seriously reduce load carrying capacity. Austenite retention, whilst desirable in moderate amounts, reduces surface hardness and compressive residual stress if present to excess. Carbide precipitation can provide sites for premature fatigue crack initiation. Localised decarburisation leads to low surface hardness and adversely modifies the favourable surface compressive stresses.

Clearly the metallurgist has an important responsibil- ity for quantifying the acceptable limits for each, particularly in the case of critical naval marine gearing. There is comprehensive book on the subject [l].

The practical ability to exercise a high degree of control over furnace atmosphere, and indeed over all aspects of the heat treatment cycle, is now becoming a reality with the advent of the oxygen probe and microprocessor based furnace instrumentation. In addition to improved atmosphere control, these facilities reduce the risks of operator error and enable automatic safety features to be incorporated.

Steel Selection

Choice of suitable steel composition is important to ensure that the required care and core hardenability are achieved at the least cost. For example, low alloy car-


SURFACE HARDENING OF MARINE GEARS BARKER/MUDD

I Figure 5. Excessive Carbide Precipitation.

burising steels will not develop the required surface the case hardness distribution including core hard- hardness or case hardness profile in large section sizes. ness (Figure 8) This arises since the metallurgical structure progression of metallurgical structure transforma-

tions during quenching formed on quenching depends on cooling rate and the the final microstructure present anywhere within the continuous cooling characteristics of the steel. These in case turn are influenced by component mass and steel com- positions respectively.

develoDed is a series of interlinked commter Drograms A valuable steel selection aid which we have Distortion and Growth

which we call STAMP - Structure and-Mateiial Prop- erties. Employing regression equations to develop C.C.T. diagrams, together with heat transfer coeffi- cients, these programs model the metallurgical changes taking place during the heat treatment process, making it possible to predict the metallurgical structures and important engineering properties produced by carburise case-hardening components of various geometries and section sizes. Figure 9 illustrates an example where an input of chemical composition, heat treatment details, carbon profile produced during carburising, and the geometric features of the component enable us to produce -

FIgure 6. Low Case Carbon Non-Martensitic Transfor- mation Products.


The major problem invariably experienced with carburise case hardening is growth, which is the result of phase changes in the material inevitable during the heat treatment, and distortion, which is non-uniform size change caused primarily by geometrical and temperature variations. For a given geometry, growth and distortion increase in severity rapidly with size.

Excessive distortion can result in -

- scrap due to the impossibility of finish machining to the required gear and gear tooth dimensions

- risk of failures in service due to excessive, perhaps local, case thinning during subsequent grinding to

Figure 7. Surface Decarburisation.

BARKERIMUDD SURFACE HARDENING OF MARINE GEARS

l lR lERIRL 6551113 6651117 C 0 . 1 2 1 0 . 1 6 4

l o 0

Figure 8. Comparison of Steel Case Hardenability.

n- 0 00 3 w 4 0 0 5 0 0 0 0 0 DISTANCE FROM SURFACE (mml

create the required tooth profile accuracy and dimensions

- excessive post heat treatment grinding costs to cor- rect the distortion. This can easily run to several thousand dollars.

The advent of skive hobbing, whereby material can be rapidly removed in the hardened state using special carbide tip hobs, can greatly reduce time and costs, but these nevertheless remain significant, and the danger of excessive case thinning is not removed.

It is vital therefore that all possible measures are taken to minimise distortion and, equally important, ensure consistent distortion behaviour. This is achieved by close attention to:

a) steel composition limits; b) prior heat treatment of the blanks; c) good furnace construction, temperature uniformity

d) suitable gear support fixture; e) uniform quenching conditions.

We can then make optimum growth and distortion allowances at the various heat treatment stages and specify grinding allowances which permit removal of all tooth distortions whilst retaining the required case depth.

Paying attention to all measures listed in (a) to (e) above, together with a close examination of their specific influence, produced substantial reductions in the distortion of gear wheels of the type illustrated in Figure 1. Overall taper was reduced from up to 0.030 in to less than 0.010 in, scatter on taper was reduced from 0.015 in to 0.005 in.

General experience is clearly invaluable when dealing with these problems, but equally important is the syste- matic collection of data on distortion trends and growth characteristics to enable quantitative predictions to be made for particular gear geometries and steel grades.

and control;

Unlike the vehicle gear field, it is important in the marine gear field to be able to predict accurately the effects of a given heat treatment, since there is seldom sufficient opportunity to utilize sacrificial gears to establish trends.

An example of the predictive approach can be seen in a paper by Bloch [2] where typical distortion tendencies for different steels and different gear dimensions are discussed together with a suggested practical approach to forecasting approximate change in diameter and length.

INDUCTION HARDENING


For naval marine gears the process employed is submerged tooth by tooth progressive hardening where the inductor passes through the space between two teeth.

The process involves inducing a high frequency cur- rent from the inductor into the near-surface regions of the tooth profile. This produces intense and extremely rapid local heating in the tooth flanks, root fillets and roots. These zones are then transformed into hard medium-carbon martensite by the quenching effects of spray jets following the inductor and the bulk oil in which the tooth being hardened is submerged.

Figure 10 shows a typical hardness profile. Whilst it is possible to produce ample case depth to suit almost any pitch of tooth, control of depth is not possible with the same degree of precision as carburise case hardening. This arises since power generator frequency has a major influence on the case depth produced and is not a process control variable. However, provided the minimum required case depth is achieved, precise control is not necessary, and we know that our 10 KHz frequency facility is entirely suitable for achieving the required case depth for a range of tooth sizes from 4 to 0.75 DP. For finer pitch teeth a higher frequency energy source is necessary to provide a shallower heating effect and avoid “back tempering” or softening of the previously hardened flank.

The residual stress profile produced is basically similar to that induced by carburise case-hardening (Figure 11) with favourable compressive stresses in the case. The balancing tensile stresses are again situated in the region of the case and core interface but typically display a higher peak value. In practice this creates no problem since the case depth is sufficient to ensure that this peak is located in a low stress region well below the tooth surface.

As with carburising, a low temperature temper is carried out after hardening to provide a tougher martensitic case and at the same time reduce the tensile residual stress levels. Resultant hardness is normally in the range 52/54Rc.

Process Control

The induction hardening machines manufactured and used at DBGI plants are shown in Figure 12. The facility



tlAT ER I AL SPECIFICATION 655R13 EN 3 6 A

CARBON 0.12 IIANGANESE 0.45

SILICON 0.20 CHROII IUfl 0.82

IIOLYBDENUfl 0.0 0 NICKEL 3.30

HEAT TREfrrflENT AUSTENISING TErTP. 935 DEG. C.

COOLED FROfl 810 DEG.C. TO 30 DEG.C.

TEnPERED AT 175 DEG.C.

4 4 nm. 100 NINS

IN orL

FOR 2.0 HRS ENGINEERING REQUIREtlENTS RULING SECTION 1.25 tlnS

E.C.D. 2.5 flnS

PROCESSING DATA CARBURISE AT DEG. C.

BOOST FOR HR/S DIFFUSE FOR HR/S

-lo 0 1 2 3 4 5

Distance frnrn surface 0 - 1 1 1

K q to histogram.

Carbide

Austrnitu

474

443

Figure 9. “Stamp Output.”



. n i o . o m , 0 3 0 .eon ,050 o60 .n7o . o m , 0 9 0 . l o o , 1 1 0 . D e p t h Below Surface (ins)

Figure 10. Typical Hardness Gradient - Induction Harden- ing.

comprises a generator, power control cabinet, gear handling machine, water cooling tank and a control console.

The inductor and spray blocks are mounted on top of the workhead transformer, which in turn is fixed to a carriage mounted on a pair of linear bearing tracks. The steel bed which carries mountings to support the gear is surrounded by a tank which can be flooded by quen- chant oil such that the carriage, inductor and tooth space to be hardened are completely immersed. The machine illustrated has been in use for over 15 years and is capable of handling gears up to approximately 150 in dia., 36 in facewidth, and over 30,000 lb in weight. The metallurgical features which need to be considered are not so numerous or vaned compared to carburise case hardening. However, control of the process parameters is in some respects more critical.

These parameters are the inductor to workpiece coupling, power to the inductor, inductor traverse speed and the oil flow. We have therefore developed standard relationships between these parameters and tooth size thus enabling us to largely dispense with test gears. (Figures 13)

Development work has also been undertaken to establish the permissible variation in a given parameter which can be tolerated without adversely affecting quality.

The inductor to workpiece gap is particularly important being of the order 1-3 mm with a relatively tight tolerance: too close and there are risks of burning or cracking; too distant and poor, patchy, or non-existant hardening will result. To ensure that the required

71 D e p t h Below S u r f a c e ( i n s )

1 I Figure 11. Typical Residual Stress Distribution-Induction Hardening.

Figure 12. Induction Hardening Installation.

tolerance is maintained across the full traverse and for each successive tooth space, well designed gear handling equipment is essential.

Equally important is the use of inductors manufactured to consistently high standards of quality and profile accuracy. The latter is achieved at DBGI by finish machining a standard block to size by numerically controlled (NC) profiling, the particulars for the NC machining being obtained from the specific gear tooth particulars to ensure complete accuracy.

Outside the establishment of these general parameters, attention needs to be paid to “corner” effects at the ends of the tooth facewidth, with the avoidance of soft spots or case thinning at one extreme and overheating or melting at the other, Figure 14. All these undesirable features can be readily overcome by adoption, as appropriate, of a number of facilities on the machine:

dwell at entry or exit of the inductor from the toothspace or both turn up or down power at entry or exit or both increase or decrease inductor traverse speed at entry or exit or both

Naval Engineers Journal, May 1984 251


I l l l l l l l l l l l l I l l l ~ 1 1 l 1 1 1 1

a m 0

u

0

Y 0

3 73

H

Tool ti 'i

Figure 13. Hardening Parameters Vs Tooth Size.

Distortion and Growth

Since bulk heating and quenching are avoided, induction hardening has the advantage of much reduced distortion compared with carburise case hardening. Any thermal distortion which does occur is primarily of consequence during the hardening process itself in terms of the need to maintain constant inductor to workpiece gap for each successive tooth space. Circularity variation of the gear after completion of hardening is not significant. As far as naval gears are concerned the gear teeth would be finished by grinding, which, because of the lower distortion, would normally be a much shorter operation than that required after carburising and hardening.

However, in the particular case of thin gear rim con- structions, problems can arise. Our research has reveal- ed that longitudinal stresses are developed across the face width during hardening, and these can result in a tendency for the rim to develop a larger diamter at the extremities compared with the centre, commonly referred to as "diabolo." This distortion increases with face width and reduced rim thickness, hence these dimensions are subject to design rules to prevent such problems arising.

NITRIDING


Nitriding differs fundamentally from the other two surface hardening processes in the physical metallurgy

i l t l n g Y100

Figure 14. End/Corner - Induction Hardening.



I ’)no 1 1

500 - I 3 0 0 1 2 7 0 1 . . . . . . , . ~

.oln ,020 . O W ,040 .050 ,060 .OIL] ,080 ,090 . l o o , 1 1 0 . i z c I 1 Depth Below S u r f a c e (ins)

Figure 15. Typical Hardness Gradient - Nitriding.

of the hardening mechanism involved. Transformation to hard martensite case is not involved. Instead, the in- wards diffusing nitrogen combines with certain alloying elements in the steel to form small nitride particles.

The high nitride case hardness is then primarily a consequence of dislocation interaction with the strain fields associated with these alloy nitrides. However, the ap- parent complexity of physical metallurgy belies a relatively simple, easily controllable process which can be carried out at relatively low temperature (below 55OoC) in the presence of a dissociated ammonia gas atmosphere or under plasma. The case depth produced, Figure 15, is dependant on actual composition and can be controlled by process time and, to a lesser extent, temperature, Due to limited nitrogen diffusion and reaction rates the achievable case depth has a maximum value of approximately 0.025 in, even with long process times. This has consequences for gear load carrying capacity which we shall examine later.

The residual stress distribution, (Figure 16) is, once again, basically similar in form to that obtained by the other processes considered, with a relatively high level of compressive stress in the case. The balancing tensile stresses tend to have peak values less than half the compressive peak. Surface hardness is primarily dependent on steel composition and to a lesser extent on prior heat treatment, specifically the tempering temperature. Naval propulsion gears have been manufactured from 3% Cr - 1% Mo steel with a surface hardness typically 850 VPN which converts to 65 Rc.

u u) 050 ,100 . 150 200 I : o

D e p t h Relow Surface ( i n s )

-60 JJ Figure 16. Typical Residual Stress Distribution - Nitriding.

Figure 17. Nitrided “White Layer.”

Process Control

Just as the process is comparatively simple, so the possibilities of metallurgical variations and the process control parameters are few.

Accurate control of temperature is important since too low a temperature will result in a shallow case, which can be corrected by re-nitriding, but too high a temperature will lead to low surface hardness which cannot be corrected.

The arguments, plasma or ion nitriding versus conventional gas nitriding are largely unimportant for marine gears. At the commercial level, plasma is quicker than gas for shallow case depths but where the maximum attainable case depth is required the differences are small. Technically, plasma offers more controllabili- ty over the depth and nature of “white layer” (Figure 17) - the shallow single or duplex E or y-iron nitride which forms on the surface and which appears white when examined under the microscope after nital et- ching.

Whilst the duplex layer formed during conventional gas nitriding its brittle and can easily exfoliate, we have no evidence that its presence in any way reduces fatigue resistance and load carrying capacity of the gear teeth. Moreover, any white layer exfoliation which might have occurred has never, to our knowledge, caused gearbox performance problems. However, should the desire for minimal white layer prevail, this can be achieved to some extent even in gas nitriding by employing infrared monitoring and microprocessor control of the percent dissociation.

Distortion and Growth

Since process temperatures are low and no quenching is involved, distortion is minimal with just a small predictable growth taking place which can usually be disregarded.



LOAD CAPACITY

The load capacity of a gear pair in any modern calculation procedure is a function of the stress permitted either at the meshing surfaces or at the tooth root. The stress at both of these locations resulting from the contact loads is a maximum at or near the surface and, therefore, if the surface layers have a higher hardness than in the core the load to produce failure will be greater than that obtained from a gear made entirely of the core material. It has long been recognized that the degree of advantage is dependant on the hardness of the case although there is some indication in the literature (Ref. 3, 4) that there seems to be a higher limit occur- ing near to 58Rc (650 Hv). Logically it can also be expected that the hardness profile below the surface layers will also affect the capacity of the gear to resist load.

There is an added effect due to the residual stress pat- tern caused by the change in volume of the hardened surface layers. As previously discussed, when the surface heat treatment process is carried out correctly this induces compressive residual stress at the surface layers which inhibits crack initiation and propagation and also enhances the load carrying capacity. There is a corollary to this however in balancing tensile residual stress at some point near the case and core junction. This tensile residual is additive to the bending tensile stress and, although not usually limiting, needs to be considered.

The hardness gradients typical of each process have been referred to previously in Figures 2, 10, and 15 and similarly typical residual stresses have been shown in Figures 3, 11, and 16. These values have been used when calculating the typical permissible stresses shown later in Table 1.

The modern calculation procedures for load capacity are based on an assessment of the stresses at the tooth flank and at the tooth root. However, they do not ade- quately distinguish the effects of the variation in surface

r w e

CH CH CH CH CH CH

N N N N N N

IH IH IH

TABLE 1. Typical Permissible Stresses. Crs Case Mod mm Ratio mm mm HV mm ZB ZC 6HLIM

80 3 .5 2 720 6.6 .88 1.0 2024 200 2 .75 5 720 20 3 4 .98 1893 200 4 .75 4 720 14 .87 1.0 2001 560 4 1.0 6 720 39 .69 1.0 1587

1250 4 1.3 8 720 88 .56 1.0 1288 1250 4 2.0 12 720 88 ,655 1.0 1507

80 3 .5 2 850 6.6 .88 1.0 2024 200 2 .5 5 850 20 .84 .82 1584 200 4 .5 4 850 14 .87 .87 1741 560 4 .5 6 850 39 .69 .80 1270

1250 4 .5 8 850 88 .56 .75 966 1250 4 .5 12 850 88 ,655 .73 1100

560 4 2.5 8 615 39 .77 .99 1677 1250 4 2.5 12 615 88 ,655 .88 1268 1250 4 4.0 12 615 88 .655 1.0 1441

= Radius of relative curvature (mm) GHLIM = Permissible surface stress (MN/mm2)


U l t i m a t e Tensile S t r e n g t h k.s.1. 200 300 a00 5 0 0 600 700 800 900

Hardness H v

Figure 18. Value of Ho.

hardness below the surface or the effect of the residual stresses. An analysis of these effects was described in a recent paper [5 ] and for a representative presentation of the comparative value of the surface hardening procedures discussed, a summary of the procedure is given below. The graphs are the result of a fundamental stress analysis and as such do not depart in principle from either the IS0 or the AGMA load capacity procedures.

Contact Permissible Stress

It can be shown that the contact permissible stress, Hlim, for a surface hardened gear can be assessed by the expression

H,, = H, X zB X z, H, is the Basic Permissible Stress 2, is the Geometry Effect Factor Z, is the Case Depth Factor

where

The values given in the following tables are of necessity simplifications. They are, however, the summary of many individual analyses and can be used for judging the safe basic stresses to be used in gear load capacity calculation.

a) Basic permissible stress (Ho). This is dependant on the hardness gradient, but for a given surface hardening process can be related to the surface hardness Hv (Figure 18). This is the stress which would be found from disc testing assuming the depth of the hardened layer had achieved the limiting value. Typical residual stress distribution for the particular surface hardening process is also included in this factor.

This accounts for the interrelationship between the surface Hertzian stress and the bending stress at the pitch line which results when the load approaches the tooth tip. The combined stress cycle can be evaluated using a Goodman diagram, and its effect can be shown to vary with the dimensionless parameter Pred/mn, where Pred is the radius of relative curvature and m, is the tooth module. This relationship also exists for through hardened gears but to a lesser degree as shown in Figure 19.

b) Geometry effects factor (ZB).

BAFtKER/MUDD SURFACE HARDENING O F MARINE GEARS

1.0 .

0.8 .

0.6

0.4

0.2 .

TABLE 2. “Analysis has indicated that typically for the surface treatments described the limiting case depth is approx-

imately proportional to tooth pitch.” Steel Limiting Case Depth

0.16 x normal module 655M13 carburised and hardened

1.0 . 0.9 .

0.8 -

722M24 nitride hardened

817M40 induction hardened

0.20 x normal module

0.32 x normal module

0.6 .

0.5 .

c) Case depth factor (Zc). If the effective case depth, defined as the depth at which hardness falls below 500 Hv, is small for a particular gear, then failure, possibly subsurface failure, may occur at some level of load. As the case depth is increased the load capacity increases up to a limiting value.

Further increases in case depth beyond this limiting value make no difference to the load capacity and the depth of case which gives rise to this limiting load capacity, known as the limiting case depth. Analysis has indicated that typically for the surface treatments here described the limiting case depth is approximately proportional to tooth pitch (Table 2). On a disc specimen it is related strictly to the radius of relative curvature.

If the actual effective case depth is less than the limiting case depth then the load which can be carried is reduced by the factor ZC (Figure 20).

LOAD CAPACITY

The simplest way of illustrating the effect of the geometry and case depth effects is to consider a range of realistic gear sizes and show how the permissible stress may vary. This is given in Table 1.

The above examples show that nitriding has a load capacity equally as high as that obtained with carburs- ing when the effective case depth of the finished gear meets the criterion of .16 x module. However, this is dif- ficult to achieve for modules larger than 2 mm and therefore for gears of larger module a lesser permissible stress is to be expected. For large gears, such as marine final ratio gears, the penalty of the thin nitrided case depth is indicated as severe.

U.T.S. - 120 k.s.i.

U.T.S. - 150 k . s . i .

- U.T.S. -2Mlk.s. i .

Surface Hardened

0.4 I o i 4 B a i i i 2 Relative Radius of Curvature/Module I Figure 19. Value of ZB.

I . * 1

I 0 1 I 0 0.2 0.4 0.6 0.8 1.0 1.2 I

I Actual Case Depth Limiting Case Depth I

Figure 20. Case Depth Factor.

Induction hardened gears however do not have a case depth limitation as size increases. The limitation for these gears tends to be due to the hardness achievable on the tooth flank, bearing in mind the possibility of back tempering. With the commonly used materials, 55 Rc (600 Hv) is feasible, and it is conceivable that some com- bination of material, quench and hardening parameters may be found which can approach 60Rc (700 Hv) satisfactorily.

BENDING STRENGTH

Contact stress in a gear increases as the square root of the load, whereas the bending stress varies as the load. Therefore surface hardened gears when loaded to their limit tend to be strength limited rather than surface limited. This can be readily controlled by selecting a larger pitch or other tooth design manipulation. The usual ACMA inscribed parabola tooth strength calculation would be used as the criterion. However, when surface hardened gears are used, the subsurface also needs consideration to take account of the varying stress pat- tern, the varying hardness and the residual stress. With carburised and hardened or with induction hardened gears there is usually no problem and adjustments may be made if necessary to the case depth, but with nitrided gears the limited case depth needs consideration. Because of the wide range of design changes which could be invoked no attempt is made here to generalize as to the relative effect of the surface treatments on bending load capacity.

CONCLUSIONS

On reflecting back over this paper you may have come to the conclusion that the possible problems to be encountered and overcome when adopting surface hardening techniques for naval marine gearing are more than a little daunting! The corollary to this, however, which I hope will be equally clear, is that the technology, plant, and expertise exist to negate these fears and enable the most progressive gear manufactur- ers to design naval marine gear units around surface hardened gearing in the knowledge that the necessary properties can be obtained and that design calculation techniques do exist to closely match these properties to permissible load carrying capacity. Continued on p. 273

26 1 Naval Engineers Journal, May 1984

MEETINGS AND SYMPOSIA

aspects of gas turbine technology, including research and development, education, systems concepts, applica- tion and operational experience. The deadline for abstracts is 1 July 1984. The abstracts should profile the paper in 50 words or less. Completed manuscripts must be received no later than 1 September 1984. To submit abstracts or for further information, contact:

Howard L. Julien, Program Chairman Raymond Kaiser Engineers Inc. Advanced Technologies Div. BB-4 P.O. Box 23210 Oakland, CA 94623-2310

Meeting on Transportation and Communications to be held 9-12 October 1984 in Genova, Italy. The theme of the meeting is telecommunications, land, maritime, air and space transportation, theory and information techniques. The papers must be typed in final form on model paper, size 21 cm/ by 30 cm/ and the final text of the papers should reach the IIC no later than 30 June 1984. Papers must not exceed eight pages; any paper ex- ceeding this limit will not be accepted. For further information contact:

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TEL: (010)294683 A call for papers has been issued by Istituto Interna- 1-16125 GENOVA, Italy

zionale delle Comunicazioni for its 32nd International

NOW HEAR THIS! ASNE Members are cordially

invited to join in the celebration at the

Commissioning Ceremonies for

USS YORKTOWN (CG-48) 4 July 1984, Yorktown, Virginia

A Trip Sponsored By The

Northern Virginia Council

Navy League of the U.S.

For Further Information Contact:

Buck Burrow, Activity Chairman Janice Delancy, Council President (202) 692-0331 (703) 845-0055

Continued from p . 261 REFERENCES

131 SAE Handbook “Fatigue Design.” Fig. 3-37. [l] Parrish G., The Influence of Microstructure on the Prop- i41 Data Handbook ESDU “Stress-& Streis,” Volume 6, No.

erties of Case Carburised Components, American Society of Metals. [5 ] Mudd, G.C. & M. France, “Evaluation of Endurance

[2] Bloch, Dr. P., Heat Treatment Distortions, Eleventh Limit” for Contact Stress in Gears, with Particular Round Table Conference on Marine Gearing, Chatham, Reference to Surface Hardened Gears,” AGMA Technical Massachusetts, 3-5 October 1977. Conference, Montreal, 17-19 October 1983.

273

73005, Royal Aeronautical Society, UK.

Naval Engineers Journal, May 1984

Documents

SURFACE HARDENED NAVAL MARINE GEARS WITH REFERENCE TO ALTERNATIVE MEANS OF SURFACE HARDENING