DNVGL-CG-0037 Calculation of crankshafts for reciprocating ......Changes - current Class guideline — DNVGL-CG-0037. Edition November 2015 Page 3 Calculation of crankshafts for reciprocating

The electronic pdf version of this document, available free of chargefrom http://www.dnvgl.com, is the officially binding version.

DNV GL AS

CLASS GUIDELINE

DNVGL-CG-0037 Edition November 2015

Calculation of crankshafts for reciprocatinginternal combustion engines

FOREWORD

DNV GL class guidelines contain methods, technical requirements, principles and acceptancecriteria related to classed objects as referred to from the rules.

© DNV GL AS November 2015

Any comments may be sent by e-mail to [email protected]

If any person suffers loss or damage which is proved to have been caused by any negligent act or omission of DNV GL, then DNV GL shallpay compensation to such person for his proved direct loss or damage. However, the compensation shall not exceed an amount equal to tentimes the fee charged for the service in question, provided that the maximum compensation shall never exceed USD 2 million.

In this provision "DNV GL" shall mean DNV GL AS, its direct and indirect owners as well as all its affiliates, subsidiaries, directors, officers,employees, agents and any other acting on behalf of DNV GL.

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Class guideline — DNVGL-CG-0037. Edition November 2015 Page 3Calculation of crankshafts for reciprocating internal combustion engines

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CHANGES – CURRENT

This is a new document.

Editorial correctionsIn addition to the above stated main changes, editorial corrections may have been made.

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CONTENTS

Changes – current...................................................................................................... 3

Section 1 General....................................................................................................... 61 Scope...................................................................................................... 62 Field of application................................................................................. 63 Principles of calculation..........................................................................64 Drawings and particulars to be submitted..............................................7

Section 2 Calculation of stresses..............................................................................111 Explanations and Basic principles.........................................................112 Calculation of stresses due to bending moments and radial forces.......113 Calculation of alternating torsional stresses.........................................17

Section 3 Evaluation of stress concentration factors................................................221 General................................................................................................. 222 Crankpin fillet....................................................................................... 243 Journal fillet (not applicable to semi-built crankshaft)......................... 254 Outlet of crankpin oil bore................................................................... 265 Outlet of journal oil bore......................................................................266 Inclined oil bores..................................................................................26

Section 4 Additional bending stresses......................................................................281 General................................................................................................. 28

Section 5 Calculation of fatigue strength................................................................. 291 General................................................................................................. 292 Empirical formulae for non-surface treated fillets.................................293 Surface treated fillets and/or oil holes.................................................32

Section 6 Acceptability criteria.................................................................................331 General................................................................................................. 332 The crankpin fillet criterion.................................................................. 333 The journal fillet criterion.....................................................................334 The oil hole criterion............................................................................ 34

Section 7 Calculation of shrink-fits of semi-built crankshaft.................................... 351 General................................................................................................. 352 Maximum permissible hole in the journal pin....................................... 36

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3 Necessary minimum oversize of shrink-fit............................................364 Maximum permissible oversize of shrink-fit......................................... 37

Appendix A Definition of stress concentration factors in crankshaft fillets............... 38

Appendix B Stress concentration factors and stress distribution at the edge of oildrillings.................................................................................................................39

Appendix C Alternative method for calculation of stress concentration factors inthe web fillet radii of crankshafts by utilising finite element method................... 40

1 General................................................................................................. 402 Model requirements..............................................................................403 Load cases............................................................................................41

Appendix D Guidance for evaluation of fatigue tests................................................471 Introduction..........................................................................................472 Evaluation of test results......................................................................473 Small specimen testing....................................................................... 534 Full size testing.................................................................................... 545 Use of existing results for similar crankshafts......................................56

Appendix E Surface tested fillets and oil bore outlets.............................................. 581 Introduction..........................................................................................582 Definition of surface treatment............................................................ 583 Calculation principles............................................................................584 Induction hardening............................................................................. 635 Nitriding................................................................................................656 Cold forming......................................................................................... 67

Appendix F Stress concentration factors in the oilbore outlets of crankshafts byutilizing finite element method.............................................................................70

1 General................................................................................................. 702 Model requirements..............................................................................703 Load cases............................................................................................71

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SECTION 1 GENERAL

1 ScopeThis class guideline contains methods for calculation of safety against fatigue failure in pin and journal filletsand in oil bores, as well as safety against slippage for semi-built crankshafts. Note that also axial bores incrankpins can be critical.In principle, any relevant calculation method may be applied for this purpose.E.g. the fatigue criteria assume that maximum resp. minimum bending stresses occur simultaneously withthe maximum resp. minimum torsional stresses. (This assumption yields results which are more or less tothe safe side.)If properly documented an evaluation of the fatigue capacity by using multi-axial fatigue methods mayreplace the criteria applied in this class guideline.A crankshaft will be approved when the calculation method as presented in this class guideline and/or theIACS UR M53 method is fulfilled. Crankshafts that cannot satisfy these rules may be subject to specialconsideration when detailed calculations and/or measurements are made to demonstrate equivalence tothese rules.

2 Field of applicationThese methods apply to engines for both main propulsion and auxiliary purpose, being designed forcontinuous operation at the specified (nominal) rating. If frequent use of overload (in excess of nominal)rating is foreseen, i.e. accumulating several millions of stress cycles, this overload should form the basisfor the calculation. Further, the methods apply to solid forged or semi-built crankshafts of forged or caststeel, and with one crank throw between main bearings. Designs with more than one crank throw betweenmain bearings, fillets between offset crankpins, fully built crankshafts, or welded crankshafts will be speciallyconsidered on basis of equivalence with the “normal” requirements.This method is also valid:

— for surface treated fillets,— when fatigue parameter influences are tested,— when working stresses are measured.

Note that surface treated fillets also include crankpin hardening that can influence the fillet fatigue strength.Principles of calculation

3 Principles of calculationThe design of crankshafts is based on an evaluation of safety against fatigue in the highly stressed areas. Thecalculation is also based on the assumption that the areas exposed to highest stresses are:

— fillet transitions between the crankpin and web as well as between the journal and web,— outlets of crankpin oil bores.— outlets of journal oil bores.

The methods of crankshaft fatigue are based on:

— A combination of bending and torsional stresses where bending fatigue strength and torsional fatiguestrength are separately assessed, i.e. also valid for anisotropic material behaviour.

— Bending stress measurements, or continuous beam analysis, or 3-point bending, depending onmanufacturer’s documentation.

— Measured stress concentration factors, or analysed with FEM as described in App.C and App.F, or by usingempirical formulae with one standard deviation above 50% reliability.

— Fatigue strength assessment based on full scale testing or small scale (see App.D), combined with acalculation procedure considering the influence of material strength, forging method, mean stress, and

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when applicable also fillet surface hardening, shot peening or cold rolling including the influence of thedepth of this surface treatment (see App.E).

For semi-built crankshafts, additional criteria cover the safety against slippage of shrink fits. In case ofrelatively small distance between the pin and the journal shrinkage diameter, the influence of the shrinkagestresses on the fatigue strength is considered.

4 Drawings and particulars to be submittedFor the calculation of crankshafts, the documents and particulars listed below shall be submitted:

— crankshaft drawing (which shall contain all data in respect of the geometrical configurations of thecrankshaft)

— type designation and kind of engine (in-line engine or V-type engine with adjacent connecting-rods,forked connecting-rod or articulated-type connecting-rod)

— operating and combustion method (2-stroke or 4-stroke cycle/direct injection, pre-combustion chamber,etc.)

— number of cylinders— rated power [kW]— rated engine speed [r/min]— direction of rotation (see Figure 1)— firing order with the respective ignition intervals and, where necessary, V-angle α v [°] (see Figure 1)

Figure 1 Designation of the cylinders

— cylinder diameter [mm]— stroke [mm]— maximum net cylinder pressure pmax [bar]— charge air pressure [bar](before inlet valves or scavenge ports, whichever applies)— mean indicated pressure [bar]— reciprocating mass per cylinder [kg]— rotating mass (ref. to crank radius) per connecting rod— rotating mass (ref. to crank radius) of cranks incl. counter weights [kg] and positions (lengthwise and

angular)— digitized gas pressure curve presented at equidistant intervals [bar versus Crank Angle] (at least every 5°

CA)— connecting-rod length LH [mm]

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— for engines with articulated-type connecting-rod (see Figure 2)

— distance to link point LA [mm]— link angle αN [°]— connecting-rod length LN [mm]

Figure 2 Articulated-type connecting-rod

— details of crankshaft material

— material designation (according to ISO, EN, DIN, AISI, etc..)— mechanical properties of material (minimum values obtained from longitudinal test specimens):

— tensile strength [N/mm²]— yield strength [N/mm²]— reduction in area at break [%]— elongation A5 [%]— impact energy – KV [J]

— type of forging(free form forged, continuous grain flow forged, drop-forged, etc; with description of the forgingprocess)

— finishing method of fillets and oil holes

— surface hardening of fillets, if applicable (induction hardened, nitrided etc.):

— extension of surface hardening (max and min)— hardness at surface (HV) (max and min)— hardness (HV) as a function of depth [mm]

— cold deformation of fillets, if applicable (shot peening, stroke peening, cold rolling):

— extension of treatment— intensity (shot peening)— depth of treatment (residual stresses) for stroke peening and cold rolling, as determined for the

applied force and roller or ball geometry.

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— particulars of alternating torsional stress calculations, see Sec.2[3].

Figure 3 Left: Crankthrow for in line engine. Figure 3 Right: Crankthrow for V- engine with 2adjacent connecting-rods

L1 = Distance between main journal centre line and crankweb centre (see also Sec.2 Figure 1 forcrankshaft without overlap)

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L2 = Distance between main journal centre line and connecting-rod centreL3 = Distance between two adjacent main journal centre lines

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SECTION 2 CALCULATION OF STRESSES

1 Explanations and Basic principlesSec.2 deals with nominal stresses that later are multiplied with stress concentration factors from Sec.3.In the case that FE calculations are made of a global model that also includes local stresses, the directlydetermined stresses shall be used in the acceptance criteria in Sec.6.In addition to the alternating stresses, the mean stresses shall be determined in order to assess the meanstress influence on the fatigue strength.The maximum and minimum bending moments and radial forces during a complete working cycle, may bedetermined by thorough analysis, e.g. FEA.Continuous beam model analysis may be applied if its relevance can be documented.If no such analyses are available, the 3-point bending may be used as described in the following.

2 Calculation of stresses due to bending moments and radial forces

2.1 AssumptionsThe calculation is based on a statically determined system, composed of a single crankthrow supported in thecentre of adjacent main journals and subjected to gas and inertia forces. The bending length is taken as thelength between the two main bearing midpoints (distance L3, see Sec.1 Figure 3 diagrams "left" and "right".The bending moment MBR, MBT are calculated in the relevant section based on triangular bending momentdiagrams due to the radial component FR and tangential component FT of the connecting-rod force,respectively (see Sec.1 3).For crankthrows with two connecting-rods acting upon one crankpin the relevant bending moments areobtained by super position of the two triangular bending moment diagrams according to phase (see Sec.1Figure 3).

2.1.1 Bending moments and radial forces acting in webThe bending moment MBRF and the radial force QRF are taken as acting in the centre of the solid web(distance L1) and are derived from the radial component of the connecting-rod force.The alternating bending- and compressive stresses due to bending moments and radial forces shall berelated to the cross-section of the crank web. The corresponding mean stresses shall also be determined.This reference section results from the web thickness W and the web width B (see Figure 1).

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Overlapped crankshaft

Crankshaft without overlap

Figure 1 Reference area of crankweb cross section

2.1.2 Bending acting in outlet of crankpin oil bore.The two relevant bending moments are taken in the crankpin cross-section through the oil bore.

Figure 2 Crankpin section through the oil bore (clockwise rotation)

MBRO = is the bending moment of the radial component of the connecting-rod force

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MBTO = is the bending moment of the tangential component of the connecting-rod force

The alternating and mean stresses due to these bending moments shall be related to the cross-sectional areaof the axially bored crankpin. Mean stresses also to be determined.

2.2 Calculation of nominal alternating bending and compressive stresses inwebThe radial and tangential forces due to gas and inertia loads acting upon the crankpin at each connecting-rodposition will be calculated over one working cycle.

Using the forces calculated over one working cycle and taking into account of the distance from the mainbearing midpoint, the time curve of the bending moments MBRF, MBRO, MBTO and the radial force QRF – asdefined in Sec.2[2.1.1] and Sec.2[2.1.2] – will then be calculated.

In the case of V-type engines, the bending moments – progressively calculated from the gas and inertiaforces – of two cylinders acting on one crankthrow are superimposed according to phase. Different designs(forked connecting-rod, articulated-type connecting-rod or adjacent connecting-rods) shall be taken intoaccount.

Where there are cranks of different geometrical configurations in one crankshaft, the calculation shall coverall crank variants.

The decisive alternating values will then be calculated according to:

where :

XN = is considered as alternating force, moment or stressXmax = is maximum value within one working cycleXmin = is minimum value within one working cycle

A) Mean stresses

The mean values XNM shall be calculated according to:

2.2.1 Nominal alternating bending and compressive stresses in web cross sectionA) Alternating stresses

The calculation of the nominal alternating bending and compressive stresses

is as follows:

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where:

σBFN [N/mm²] = nominal alternating bending stress related to the webMBRFN [Nm] = alternating bending moment related to the centre of the web (see Sec.1 Figure 3)

Weqw [mm3] = first moment of area (sectional modulus) related to cross-section of web

Ke = empirical factor considering to some extent the influence of adjacent crank and bearing restraint with:

Ke = 0.8 for crosshead enginesKe = 1.0 for trunk engines

σQFN [N/mm²] = nominal alternating compressive stress due to radial force related to the web

F [mm²] = area related to cross-section of web

F = B · W

B) Mean stresses

The mean values shall be calculated according to:

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2.2.2 Nominal Bending stress in outlet of crankpin oil boreA) Alternating stresses

The calculation of nominal alternating bending stress is as follows:

where:

σBON [N/mm²]= nominal alternating bending stress related to the crankpin diameterMBON [Nm] = alternating bending moment calculated at the outlet of crankpin oil bore

with

and ψ [°] angular position (see Figure 2)

We(red) [mm3] first moment of area (sectional modulus) related to cross-section of axially bored crankpin

B) Mean stresses

The mean values shall be calculated according to:

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2.2.3 Nominal bending stresses in outlet of journal oil boreA) Alternating stressIf the 3-point bending method is applied (resulting in zero bending moment), the alternating bendingmoment in the journal MBONG shall be taken at least as 0.5 · MBON. The nominal alternating bending stress isthen:

B) Mean stressIf the 3-point bending method is applied, the mean bending moment in the journal MBONGM may bedisregarded.

2.3 Calculation of bending stresses in filletsA) Alternating stresses

The calculation of stresses shall be carried out for the crankpin fillet as well as for the journal fillet.

For the crankpin fillet:

σBH = ± (αB · σBFN)

where:

σ BH [N/mm2] = alternating bending stress in crankpin filletα B [-] = stress concentration factor for bending in crankpin fillet (determination – see Sec.3).

For the journal fillet (not applicable to semi-built crankshaft):

where:

σ BG [N/mm2] = alternating bending stress in journal filletβ B [-] = stress concentration factor for bending in journal fillet (determination - see Sec.3)

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β Q [-] = stress concentration factor for compression due to radial force in journal filletdetermination – see Sec.3)

B) Mean stresses

The mean stresses shall be calculated according to:


σBHM = αB · σBFNM

For the journal fillet:

Note that the signs above are only correct when the bending moment due to max gas pressure is more thantwice the bending moment due to mass forces in TDC.

2.4 Calculation of bending stresses in outlet of crankpin oil boreA) Alternating stress

σBO = ± ( γB · σBON)

where:

σ BO [N/mm2] = alternating bending stress in outlet of crankpin oil boreγ B [-] = stress concentration factor for bending in crankpin oil bore (determination - see Sec.3)

B) Mean stress

The mean stress shall be calculated according to:

σBOM = γB · σBONM

2.5 Calculation of bending stresses in outlet of journal oil boreA) Alternating stress

σBOG = ± ( γBG · σBONG)

where:

σBONG [N/mm2] = alternating bending stress in outlet of journal oil boreγBG [-] = stress concentration factor for bending in journal oil bore (determination – see Sec.3)

B) Mean stress

σBOGM = γBG · σBONGM

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3 Calculation of alternating torsional stresses

3.1 GeneralThe calculation for nominal alternating torsional stresses shall be undertaken by the engine manufactureraccording to the information contained in Sec.2 [3.2].The manufacturer shall specify the maximum nominal alternating torsional stress.

3.2 Calculation of nominal alternating torsional stressesThe max. and min. torques shall be ascertained for every mass point of the complete dynamic system andfor the entire speed range by means of a harmonic synthesis of the forced vibrations from the 1st order upto and including the 15th order for 2-stroke cycle engines and from the 0.5th order up to and including the12th order for 4-stroke cycle engines. Whilst doing so, allowance shall be made for the damping that exists inthe system and for unfavourable conditions (misfiring *) in one of the cylinders). The speed step calculationshall be selected in such a way that any resonance found in the operational speed range of the engine shallbe detected.

*) Misfiring is defined as a cylinder condition when no combustion occurs but only compression cycle.

Where barred speed ranges are necessary, they shall be arranged so that satisfactory operation is possibledespite their existence. There shall be no barred speed ranges above a speed ratio of ℓ ≥ 0.8 for normal firingconditions.

The values received from such calculation shall be submitted. Alternatively the crankshaft assessment can bebased on the specified maximum alternating torsional stress.

For type approvals it is advised to include a variation of the various standard dampers respectively flywheelsetc. Directly coupled engines should take typical installations into account. Flexibly coupled engines maybe considered as “cut” in the elastic coupling. Where calculations are not available as e.g. for some typeapprovals, the manufacturer shall state a limiting vibratory torque to be used in the calculation.

The nominal alternating torsional stress in every mass point, which is essential to the assessment, resultsfrom the following equation:

where:

τN [N/mm²] = nominal alternating torsional stress referred to crankpin or journal

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MTN [Nm] = maximum alternating torqueWP [mm3] = polar first u of area (sectional modulus) related to cross-section of axially bored crankpin

or bored journalMTmax = maximum value of the torque.MTmin = minimum value of the torque

For the purpose of the crankshaft assessment, the nominal alternating torsional stress considered in furthercalculations is the highest calculated value, according to above method, occurring at the most torsionallyloaded mass point of the crankshaft system.

Where barred speed ranges exist, the torsional stresses within these ranges shall not be considered forassessment calculations.

The approval of crankshaft will be based on the installation having the largest nominal alternating torsionalstress (but not exceeding the maximum figure specified by engine manufacturer).

Thus, for each installation, it shall be ensured by suitable calculation that this approved nominal alternatingtorsional stress is not exceeded. This calculation shall be submitted for assessment.

The nominal mean torsional stress is:

This may be determined in every mass point which is essential to the crankshaft assessment, but a morepractical approach shall use the mean stress at the crank that has the highest alternating torque. This canbe determined from a quasistatic torque evaluation through the crankshaft (vectorial torque summationassuming clamped end and no dynamics).

3.3 Calculation of torsional stresses in fillets and outlet of oil boresThe calculation of stresses shall be carried out for the crankpin fillet, the journal fillet and the outlet of the oilbores.


A) Alternating stress

where :

τH [N/mm2] = alternating torsional stress in crankpin filletσT [-] = stress concentration factor for torsion in crankpin fillet (determination - see Sec.3)

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τN [N/mm2] = nominal alternating torsional stress related to crankpin diameter

B) Mean stress

where:

τNM[N/mm2] = nominal mean torsional stress related to crankpin diameter

For the journal fillet (not applicable to semi-built crankshafts):


where:

τG [N/mm2] = alternating torsional stress in journal filletβT [-] = stress concentration factor for torsion in journal fillet (determination - see Sec.3)τN [N/mm2] = nominal alternating torsional stress related to journal diameter

B) Mean stress

where:

τNM [N/mm2] = nominal mean torsional stress related to journal diameter.

For the outlet of crankpin oil bore:


where :

σTO [N/mm2] = alternating stress in outlet of crankpin oil bore due to torsionγT [-] = stress concentration factor for torsion in outlet of crankpin oil bore (determination - see

Sec.3)τ

N [N/mm2] = nominal alternating torsional stress related to crankpin diameter

B) Mean stress

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where

τNM [N/mm2] = nominal mean torsional stress related to crankpin diameter.

For the outlet of journal oil bore


where

σTGO [N/mm2] = alternating stress in outlet of journal oil bore due to torsionγTG [-] = stress concentration factor for torsion in outlet of journal oil bore (see Sec.3)τN [N/mm2] = nominal alternating torsional stress related to journal diameter

B) Mean stress

where

τNM [N/mm2] = nominal mean torsional stress related to journal diameter.

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SECTION 3 EVALUATION OF STRESS CONCENTRATION FACTORS

1 GeneralThe stress concentration factors are evaluated by means of the formulae according to [2], [3] and [4]applicable to the fillets and crankpin oil bore of solid forged web-type crankshafts and to the crankpin filletsof semi-built crankshafts only. It shall be noticed that stress concentration factor formulae concerning theoil bore are only applicable to a radially drilled oil hole. All formulae are based on investigations of FVV(Forschungsvereinigung Verbrennungskraftmaschinen) for fillets and on investigations of ESDU (EngineeringScience Design Unit) for oil holes. All crank dimensions necessary for the calculation of stress concentrationfactors are shown in Figure 1.

The stress concentration factors in bending (αB, βB) are defined as the ratio between the maximum absolutevalue of the PRINCIPAL stresses in the fillets and the absolute value of the nominal bending stress referred tothe web cross section.

The stress concentration factor for compressive stresses in the journal fillet (βQ) is defined as the ratiobetween the maximum fillet stress (absolute value) and the nominal compressive stress (absolute value)related to the web cross section.

The stress concentration factor for torsion (αT, βT) is defined as the ratio of the maximum equivalent shearstress – occurring in the fillets under torsional load – to the nominal torsional stress related to the axiallybored crankpin or journal cross-section (see App.A).

The stress concentration factors for bending (γB) and torsion (γT) are defined as the ratio of the maximumprincipal stress – occurring at the outlet of the crankpin oil-hole under bending and torsional loads – to thecorresponding nominal stress related to the axially bored crankpin cross section (see App.B). For outlet boresin journals the same definitions apply.

When reliable measurements and/or calculations are available, which can allow direct assessment of stressconcentration factors, the relevant documents and their analysis method shall be submitted in order todemonstrate their equivalence to present rules evaluation.

This is always to be performed when dimensions are outside of any of the validity ranges for the empiricalformulae presented in [2].

App.C describes how FE analyses can be used for this purpose. Care should be taken to avoid mixingequivalent (von Mises) stresses and principal stresses.

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Figure 1 Crank dimensions

Actual dimensions:

D [mm] = crankpin diameterDBH [mm] = diameter of axial bore in crankpinDO [mm] = diameter of oil bore in crankpinDOG [mm] = diameter of oil bore in journalRH [mm] = fillet radius of crankpinTH [mm] = recess of crankpin filletDG[mm] = journal diameterDBG[mm] = diameter of axial bore in journalRG [mm] = fillet radius of journalTG [mm] = recess of journal filletE [mm] = pin eccentricityS [mm] = pin overlap

W (*)

[mm]= web thickness

B (*)

[mm]= web width

(*) In the case of 2 stroke semi-built crankshafts:

— when TH > RH, the web thickness shall be considered as equal to:— Wred = W – (TH – RH) (refer to Sec.2 Figure 1)— web width B shall be taken «in way of» crankpin fillet radius centre according to Sec.2 Figure 1

The following related dimensions will be applied for the calculation of stress concentration factors in :

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Crankpin fillet Journal fillet

r = RH / D rG = RG / DG

s = S/D sG = S/DG

w = W/D (with overlap) wG = W/DG (with overlap)

w = Wred/D (without overlap) wG = Wred/DG (without overlap)

b = B/D bG = B/DG

dO = DO/D dOG = DOG/DG

dH = DBH/D dG =DBG/DG

tH = TH/D tH = TH/D

tG = TG/D tG = TG/D

Stress concentration factors are valid for the ranges of related dimensions for which the investigations havebeen carried out. (It should be noted that the FVV investigations were made with models having D=DG. Usingthe empirical formulae on crankshafts with DG>D may lead to journal fillet stress concentration factors thatare more uncertain than when D=DG.)

Ranges are as follows:

s ≤ 0.5

0.2 ≤ w ≤ 0.8

1.1 ≤ b ≤ 2.2

0.03 ≤ r ≤ 0.13

0 ≤ dG ≤ 0.8

0 ≤ dH ≤ 0.8

0 ≤ dO ≤ 0.2

Low range of s can be extended down to large negative values provided that:

— if calculated f (recess) < 1 then the factor f (recess) shall not be considered (f (recess) = 1)— if s < - 0.5 then f (s,w) and f (r,s) shall be evaluated replacing actual value of s by - 0.5.

The same ranges of validity apply when the related dimensions are referred to the journal diameter.

As the original (FVV) empirical formulae for the crankpin and the journal fillets were established on the basisof 50 percent reliability, and have standard deviations of approximately 10 percent (approximately 7 percentof torsion), the coefficients have been increased by a factor for bending (presently 1,10) and a factor fortorsion (presently 1,07) for the purpose of higher reliability.

2 Crankpin filletThe stress concentration factor for bending (αB) is:

αB = KαB · f (s, w) · f (w) · f (b) · f (r) · f (dG) · f (dH) · f (recess)

where:

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KαB = 3.30 (applicable to principal stresses and one standard deviation above 50% reliability)f (s,w) = -4.1883 + 29.2004 · w – 77.5925 · w² + 91.9454 · w3 – 40.0416 · w4 + (1-s) · (9.5440 –

58.3480 · w + 159.3415 · w² - 192.5846 · w3 + 85.2916 · w4) + (1-s)² · (-3.8399 + 25.0444· w – 70.5571 · w² + 87.0328 · w3 – 39.1832 · w4)

f(w) = 2.1790 · w0.7171

f(b) = 0.6840 – 0.0077 · b + 0.1473 · b2

f(r) = 0.2081 · r(-0.5231)

f(dG) = 0.9993 + 0.27 · dG – 1.0211 · dG² + 0.5306 · dG3

f (dH) = 0.9978 + 0.3145 · dH – 1.5241 · dH² + 2.4147 · dH3

f(recess) = 1 + (tH + tG) · (1.8 + 3.2 · s)

The stress concentration factor for torsion (αT) is:

αT = KαT · f (r,s) · f (b) · f (w)

where:

KαT = 0.856 (applicable to one standard deviation above 50% reliability).f (r,s) = r (-0.322 + 0.1015 · (1-s))

f (b) = 7.8955 – 10.654 · b + 5.3482 · b² - 0.857 · b3

f (w) = w (-0.145)

3 Journal fillet (not applicable to semi-built crankshaft)The stress concentration factor for bending (βB) is:

βB = 3.33 · fB (sG,wG) · fB (w) · fB (b) · fB (r) · fB (dG) · fB (dH) · f (recess)

(applicable to principal stresses and one standard deviation above 50% reliability).

where:

fB (sG,wG) = –1.7625 + 2.9821 · wG – 1.5276 · wG² + (1 – sG) · (5.1169 – 5.8089 · wG + 3.1391 · wG²) +(1 – sG)² · (-2.1567 + 2.3297 · wG – 1.2952 · wG²)

fB (wG) = 2.2422 · wG 0.7548

fB (b) = 0.5616 + 0.1197 · bG + 0.1176 · bG2

fB (rG) = 0.1908 · rG (-0.5568)

fB (dG) = 1.0012 – 0.6441 · dG + 1.2265 · dG2

fB (dH) = 1.0022 – 0.1903 · dH + 0.0073 · dH2

f (recess) = 1 + (tH + tG) · (1.8 + 3.2 · s)

The stress concentration factor for compression (βQ) due to the radial force is:

βQ = 3.69 · fQ (s) · fQ (w) · fQ (b) · fQ (r) · fQ (dH) · f (recess)

(applicable to principal stresses and one standard deviation above 50% reliability.)

where:

fQ (sG) = 0.4368 + 2.1630 · (1-sG) – 1.5212 · (1-sG)²

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fQ (wG) =

fQ (bG) = - 0.5 + bG

fQ (rG) = 0.5331 · rG(- 0.2038)

fQ (dH) = 0.9937 – 1.1949 · dH + 1.7373 · dH²

f (recess) = 1 + (tH + tG) · (1.8 + 3.2 · s)

The stress concentration factor for torsion (βT) is:

βT = 0.856 · f (rG,sG) · f (bG) · f (wG)

(applicable to one standard deviation above 50% reliability.)

where:

f (rG,sG) = rG (-0.322 + 0.1015 · (1-s))

f (bG) = 7.8955 – 10.654 · bG + 5.3482 · bG² - 0.857 · bG3

f (wG) = wG (-0.145)

4 Outlet of crankpin oil boreThe stress concentration factor for bending (γB) is:

The stress concentration factor for torsion (γT) is:

5 Outlet of journal oil boreThe stress concentration factor for bending (γBG) is:

The stress concentration factor for torsion (γTG) is:

6 Inclined oil boresInclined oil bores have higher stress concentrations than oil bores perpendicular to the surface.

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For small deviations from perpendicular oil bore exit the stress concentration factors in [4] and [5] shall bemultiplied with:

For bending:

For torsion:

Range of validity:

φ < 25 degrees

ψ < 40 degrees

where

φ = is the angular deviation (degrees) from perpendicular oil bore direction as projected in thetransverse (crosswise) plane.

ψ = is the angular deviation (degrees) from perpendicular oil bore direction as projected longitudinally inthe crankpin plane.

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SECTION 4 ADDITIONAL BENDING STRESSES

1 GeneralIn addition to the alternating bending stresses in fillets (see Sec.2 [2.3]) further bending stresses due tomisalignment and bedplate deformation as well as due to axial and bending vibrations shall be considered byapplying σadd as given by table:

Type of engine σadd [N/mm2]

Crosshead engines ± 30 (*)

Trunk piston engines ± 10

(*) The additional stress of ± 30 N/mm2 is composed of two components:

1) an additional stress of ± 20 N/mm2 resulting from axial vibrations2) an additional stress of ± 10 N/mm2 resulting from misalignment/bedplate deformation.

It is recommended that a value of ± 20 N/mm2 is used for the axial vibration component for assessmentpurposes where axial vibration results of the complete dynamic system (engine/shafting/gearing/propeller)are not available.

Where axial vibration calculation results of the complete system are available, the calculated figures may beused instead.

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SECTION 5 CALCULATION OF FATIGUE STRENGTH

1 GeneralThe fatigue strength is defined as the permanently endurable (i.e. >>107 cycles) alternating principal stressconsidering influences as mean stress, mechanical strength, size effect, stress gradient, production method,surface roughness, and surface treatment as hardening and/or cold deformation if applicable.The fatigue strength may be determined by tests or by empirical formulae as given below. Test resultsmay also be combined with pertinent elements of the empirical formulae; e.g. push-pull test results can becombined with mean stress influence, and so on.When the fatigue strength is determined by means of testing (full size or small specimen), the number of testand the statistical evaluation method shall be selected so as to enable determination of the fatigue strengthas the average value minus one standard deviation. For more details, see App.D.It is also essential that the tested material is as representative as possible for the future production; i.e.material cleanliness, production method, surface roughness, and mechanical strength shall not be superior tothe minimum requirements for the production.When full size tests are used to determine fatigue strength of crankshafts with surface treated fillets (and/or oil holes), the acceptance shall be combined with an approved procedure for quality control of the surfacetreatment.

2 Empirical formulae for non-surface treated filletsThe fatigue strengths in bending and torsion for web fillets are determined as:

For oil holes:

— for Kb, Kt and Kh, see [2.1]— for , see [2.2]

—— for f(Ra), see [2.3]— for f(χ), see [2.4]— for f(σM), see [2.5]— for f(τM), see [2.5].

When crankpins and/or journals are surface hardened, special attention shall be given to the transition zonesbetween surface hardened areas and the soft core material, see [3.1]. See also App.E.

2.1 Influence of production methods.Kb and Kt represent the influence of various production methods on fatigue in web fillet, and Kh in oil holes

Groups Kb Kt Kh

Group 1: continuous grain flow forged 1) 1.1 1.05 1

Group 2: free form forged 2) 1 1 1

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Groups Kb Kt Kh

Group 3: slab forged 3), form cut and twisted 0.9 0.9 1

Group 4: cast steel 0.7 0.85 0,7

Group 5: super clean 4) forged steel see below 4)

1) Continuous grain flow (cgf) forged means documented forging that results in a flowof segregations tangential to the main principal stress direction in the web fillets.It does not apply to oil holes (where Kh = 1). The documentation shall be madeon a crankthrow that is cut in the crank plane and etched in order to illustrate thegrain flow. It is essential that this is made with fillets in finished condition. I.e. forrecessed fillets, cgf may be difficult to obtain as the recess normally cuts the grainflow.

2) Free form forging means all kinds of forging processes that are neither cgf nor slabforging (see footnote 3).

3) Slab forging means processes where a forged slab is oxycut or machined to formcrankthrows, and then twisted into positions. All similar processes where materialfrom the centre of the slab ends up in the web fillets fall into this category. Thisdoes not apply to oil holes provided that they are not in a part that origins from theslab centre i.e. normally Kh = 1 for oil holes.

4) For steels with a especially high cleanliness (see; DNV GL rules for classification:Ships DNVGL-RU-SHIP Pt.2 Ch.2 Sec.5) the factors Kb = Kt = Kh = 1.1 for webfillets and oil holes and apply for both cgf and free form forging. When super-cleansteels are combined with slab forging,

5) Kb = Kt = 1.0 for web fillets and Kh = 1.1 for oil holes.

2.2 Mechanical strength influence.f(σB;σ0.2) represents the influence of the mechanical strength on the push-pull fatigue strength.f(σB;σ0.2) is the smaller value of

0.30 ∙ σB + 40and

0.40 ∙ σ0.2 + 70(When using these two formulae realistic values shall be used. Generalised material minimum values fromclassification rules tend to give wrong relations between σ0.2 and σB.)It should be noted that use of σB>800 MPa or σ0.2 > 600 MPa may require stricter requirements forcleanliness than those given in some material specifications.

2.3 Surface roughness influence.f(Ra) represents the influences of surfaces roughness.The surface roughness condition of all web fillets and oil holes (to a depth of at least 1,5 times the oil borediameter) is assumed to be smooth grinding or better, preferably polishing. Web fillets should be polishedalong the fillet contour. Hence the f(Ra) is unity. However, for oil bores that are not smoothly finished to theabove mentioned depth, a factor of f(Ra) = 0.9 applies.

2.4 Stress gradient influence.f(χ) represents the size influence due to the stress gradient.The other size influences such as statistical and metallurgical are not included because:

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— the statistical is covered by the use of one standard deviation below average for test results, respectivelycovered by the empirical formulae for push-pull

— the metallurgical is covered by use of representative test specimens taken from one (or both) extendedend(s) of the forged crankshafts or equivalent for single throws.

f(χ) is different for bending and torsion and may be taken as:

for bending

for torsion

where

RX is RH for crankpin fillets and RG for journal fillets and for oil bores.

However, Rx shall not be taken <2mm for calculation purposes.

2.5 Mean stress influenceThe mean stress influence is considered by means of a parabolic relation as:

respectively

where

For crankpin fillets: σMean = σBHM, see Sec.2 [1.3] B

τMean = τHM, see Sec.2 [2.3] B

For journal fillets: σMean = σBGM, see Sec.2 [1.3] B

τMean= τGM, see Sec.2 [2.3] B

It should be noted that crankpin fillet mean bending stresses normally are positive (tension) and hencereduce the fatigue strength. Journal fillet mean bending stresses are normally negative (compression) andhence increase the fatigue strength. For torsion, any mean stress regardless direction reduces the fatiguestrength.For oil holes in crankpin:

For σBOM, see Sec.2 [2.4] BFor σTOM, see Sec.2 [3.3] B.For oil holes in journals:

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— for σBOGM, see Sec.2 [2.5] B— for σTOGM, see Sec.2 [3.3] B

3 Surface treated fillets and/or oil holes

3.1 GeneralSurface treatment of fillets and oil holes means any method aiming at increased fatigue strength in localareas. It is divided into surface hardening and cold deformation.Surface hardening means all kinds of thermal treatment such as induction hardening and nitriding. Theseprocesses result in increased surface hardness, and if properly made, also compressive residual stresses to acertain depth. However, especially for induction hardening, the transition zones depth-wise and at the end ofthe layer have tensile residual stresses. It is therefore essential to check also transition zones for fatigue.Cold deformation means all kinds of “cold” treatments such as shot peening, stroke peening, cold rollingand ball coining (oil holes). These processes result in negligible hardness increase but in considerablecompressive residual stresses to a certain depth. Also with these processes the transition zones shall bechecked for fatigue.See App.D for fatigue testing and App.E for evaluation of safety against fatigue in surface treated fillets and/or oil bore outlets.

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SECTION 6 ACCEPTABILITY CRITERIA

1 GeneralFor the fatigue criteria mentioned in [2], [3] and [4] a common minimum safety factor applies, as:

S = 1.15All stresses are in N/mm2.

2 The crankpin fillet criterionThe stresses in the crankpin fillet shall fulfil the following criterion:

where

σH = σBH + σadd, see Sec.2 [2.3] respectively Sec.4.τH = see Sec.2 [3.3].σfH = σf as determined in Sec.5 for the crankpin filletτfH = τf as determined in Sec.5 for the crankpin fillet

For semi-built crankshafts with a relatively low distance between the pin and the shrunk in journal

(s = S / D > –0.1), it is necessary to consider the influence of the shrinkage stresses on the crankpin filletfatigue strength. This applies in particular to the approximately 45˚ crankpin plane.

The above mentioned criterion may be used, however, with some modifications as:

— nominal bending stresses as well as stress concentration factors to be calculated in the 45˚ plane.

Shrinkage stresses in the fillet of the 45˚ plane to be added to the mean bending stresses when determiningthe fatigue strength.

3 The journal fillet criterion(Not applicable to semi-built crankshafts.)

The stresses in the journal fillet shall fulfil the following criterion:

where

σG = σBG + σadd, see Sec.2 [2.3] respectively Sec.4.τG = see Sec.2 [3.3].

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σfG = σf as determined in Sec.5 for the journal filletτfG = τf as determined in Sec.5 for the journal fillet

4 The oil hole criterionThe stresses in the oil holes shall fulfil the following criteria:

Outlet of crankpin bores

where

σBO = see Sec.2 [2.4] A.σTO = see Sec.2 [2.2].σfO = σfh as determined in Sec.5 for the crankpin oil hole, with attention to Sec.6 [3] if the crankpin is

surface hardened.

Outlet of journal bores

where

σBOG = see Sec.2 [2.5] A.σTOG = see Sec.2 [3.3].σfOG = σfh as determined in Sec.5 for the journal oil hole, with attention to Sec.5 [3.1] if the journal is

surface hardened.

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SECTION 7 CALCULATION OF SHRINK-FITS OF SEMI-BUILTCRANKSHAFT

1 GeneralThe criteria given in [2] to [4] serve the purpose of both safe torque transmission and to prevent detrimentalfretting under the edge of the shrinkage.

All crank dimensions necessary for the calculation of the shrink-fit are shown in Figure 1.

Figure 1 Crankthrow of semi-built crankshaft

Where:

DA [mm] = outside diameter of web or twice the minimum distance x between centre-line of journals andouter contour of web, whichever is less

DS [mm] = shrink diameterDG [mm] = journal diameterDBG [mm] = diameter of axial bore in journalLS [mm] = length of shrink-fitRG [mm] = fillet radius of journaly [mm] = distance between the adjacent generating lines of journal and pin y ≥ 0.05 · DS

Where y is less than 0.1 · DS special consideration shall be given to the effect of the stressdue to the shrink-fit on the fatigue strength at the crankpin fillet.

Respecting the radius of the transition from the journal to the shrink diameter, the following should becomplied with:

RG ≥ 0.015 · DG

and

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RG ≥ 0.5 · (DS – DG)

where the greater value shall be considered.

The actual oversize Z of the shrink-fit shall be within the limits Zmin and Zmax calculated in accordance with[3] and [4].

In the case where condition in [2] cannot be fulfilled then methods in [3] and [4] calculation of Zmin andZmax are not applicable due to multizone-plasticity problems. In such case Zmin and Zmax shall be establishedbased on FEM calculations.

2 Maximum permissible hole in the journal pinThe maximum permissible hole diameter in the journal pin is calculated in accordance with the followingformula:

where :

SR [-] = safety factor against slipping, however a value not less than 2 shall be taken unlessdocumented by experiments.

Mmax [Nm] = absolute maximum value of the torque MTmax in accordance with Sec.2 [3.2]μ [-] = coefficient for static friction, however a value not greater than 0.2 shall be taken unless

documented by experiments.σSP [N/mm2] = minimum yield strength of material for journal pin

This condition serves to avoid plasticity in the hole of the journal pin.

3 Necessary minimum oversize of shrink-fitThe necessary minimum oversize is determined by the greater value calculated according to:

and

where:

Zmin [mm] = minimum oversize

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Em [N/mm²] = Young’s modulusσSW [N/mm²] = minimum yield strength of material for crank web

QA [-] = web ratio,

QS [-] = shaft ratio,

4 Maximum permissible oversize of shrink-fitThe maximum permissible oversize is calculated according to:

Higher values may be accepted for forged steel, but limited to the shrinkage amount that result in aplasticized zone of 30% of the minimum web thickness in radial direction.

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APPENDIX A DEFINITION OF STRESS CONCENTRATION FACTORS INCRANKSHAFT FILLETS

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APPENDIX B STRESS CONCENTRATION FACTORS AND STRESSDISTRIBUTION AT THE EDGE OF OIL DRILLINGS

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APPENDIX C ALTERNATIVE METHOD FOR CALCULATION OFSTRESS CONCENTRATION FACTORS IN THE WEB FILLET RADII OFCRANKSHAFTS BY UTILISING FINITE ELEMENT METHOD

1 GeneralThe objective of the analysis is to substitute the analytically calculated stress concentration factors (SCF) atthe crankshaft fillets by suitable FEM calculated figures. The analytical method is based on empirical formulaedeveloped from strain gauge measurements of various crank geometries. Use of these formulae beyondany of the various validity ranges can lead to erroneous results in either direction, i.e. results that moreinaccurate than indicated by the mentioned standard deviations. Therefore the FEM-based method is highlyrecommended.The SCF’s calculated according to the rules of this appendix are defined as the ratio of stresses calculated byFEM to nominal stresses in both journal and pin fillets. When used in connection with the present method inM53 von Mises stresses shall be calculated for bending and principal stresses for torsion or when alternativemethods are considered.The procedure as well as evaluation guidelines are valid for both solid cranks and semi-built cranks (exceptjournal fillets).The analysis shall be conducted as linear elastic FE analysis, and unit loads of appropriate magnitude appliedfor all load cases.The calculation of SCF at the oil bores is at present not covered by this document.It is advised to check the element accuracy of the FE solver in use, e.g. by modelling of a simple geometryand compare the stresses obtained by FEM with the analytical solution for pure bending and torsion.Boundary element method (BEM) may be used instead of FEM.

2 Model requirementsThe basic recommendations and perceptions for building the FE-model are presented in App.C [2.1]. It isobligatory for the final FE-model to fulfil the requirement in App.C [2.3].

2.1 Element mesh recommendationsIn order to fulfil the mesh quality criteria it is advised to construct the FE-model for the evaluation of stressconcentration factors according to the following recommendations:

— the model consists of one complete crank, from main bearing centre-line to opposite side main bearingcentre-line

— element types used in the vicinity of the fillets:

— 10 node tetrahedral elements— 8 node hexahedral elements— 20 node hexahedral elements

— mesh properties in fillet radii. The following applies to ±90 degrees in circumferential direction from thecrank plane:

— maximum element size a = r/4 through the entire fillet as well as the in circumferential direction.When using 20 node hexahedral elements, the element size in the circumferential direction may beextended up to 5 · a. In the case of multi-radii fillet, r is the local fillet radius. (If 8 node hexahedralelements are used even smaller element size is required to meet the quality criteria.)

— recommended manner for element size in fillet depth direction

— first layer thickness equal to element size of a— second layer thickness equal element to size of 2 · a

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— third layer thickness equal element to size of 3 · a

— minimum 6 elements across web thickness.

— generally the rest of the crank should be suitable for numeric stability of the solver

— counterweights only shall be modelled when influencing the global stiffness of the crank significantly

— modelling of oil drillings not necessary as long as the influence on global stiffness is negligible or theproximity to the fillet is more than 2r, Figure 1 below

— drillings and holes for weight reduction shall be modelled

Figure 1 Proximity to fillet

2.2 MaterialUR M53 does not consider material properties such as Young’s Modulus (E) and Poisson’s ratio (n). In FEanalysis those material parameters are required, as strain is primarily calculated and stress is derived fromstrain using the Young’s Modulus and Poisson’s ratio. Reliable values for material parameters shall be used,either as quoted in literature or measured on representative material samples.

For steel the following is advised: E= 2.05∙105 MPa and ν=0.3

2.3 Element mesh quality criteriaIf the actual element mesh does not fulfil any of the following criteria at the examined area for SCFevaluation, then a second calculation with a refined mesh shall be performed.

2.3.1 Principal stresses criterionThe quality of the mesh should be assured by checking the stress component normal to the surface of thefillet radius. Ideally, this stress should be zero. With principal stresses σ1, σ2 and σ3 the following criterion isrequired:

min(|σ1|,|σ2|,|σ3|) < 0.03 · max(|σ1|,|σ2|,|σ3|)

2.3.2 Averaged/un-averaged stresses criterionThe criterion is based on observing the discontinuity of stress results over elements at the fillet for thecalculated SCF:Un-averaged nodal stress results calculated from each element connected to a nodei should differ less than5% from the 100% averaged nodal stress results at this nodei at the examined location.

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3 Load casesTo substitute the analytically determined SCF in UR M53 the following load cases shall be calculated.

3.1 TorsionIn analogy to the testing apparatus used for the investigations made by FVV the structure is loaded in puretorsion. In the model surface warp at the end faces is suppressed.Torque is applied to the central node located at the crankshaft axis. This node acts as master node with 6degrees of freedom and is connected rigidly to all nodes of the end face.Boundary and load conditions are valid for both in-line and V- type engines.

Figure 2 Boundary and load conditions for the torsion case

For all nodes in both journal and crank pin fillet principal stresses are extracted and the equivalent torsionalstress is calculated:

The maximum value taken for the subsequent calculation of the SCF:

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Where tN is nominal torsion stress referred to the crankpin respectively journal as per UR M53 2.2.2 with thetorsional torque τ:

3.2 Pure bending (4 point bending)In analogy to the testing apparatus used for the investigations made by FVV the structure is loaded in purebending. In the model surface warp at the end faces is suppressed.The bending moment is applied to the central node located at the crankshaft axis. This node acts as masternode with 6 degrees of freedom and is connected rigidly to all nodes of the end face.Boundary and load conditions are valid for both in-line and V- type engines.For all nodes in both journal and pin fillet von Mises equivalent stresses σequiv are extracted. The maximumvalue is used to calculate the SCF according to:

Nominal stress sN is calculated as per UR M53 2.1.2.1 with the bending moment M:

3.3 Bending with shear force (3 point bending)This load case is calculated to determine the SCF for pure transverse force (radial force, βQ) for the journalfillet.In analogy to the testing apparatus used for the investigations made by FVV the structure is loaded in 3-pointbending. In the model, surface warp at the both end faces is suppressed. All nodes are connected rigidly tothe centre node; boundary conditions are applied to the centre nodes. These nodes act as master nodes with6 degrees of freedom.The force is applied to the central node located at pin centre-line of the connecting rod. This node isconnected to all nodes of the pin cross sectional area. Warping of the sectional area is not suppressed.Boundary and load conditions are valid for in-line and V- type engines. V-type engines can be modelled withone connecting rod force only. Using two connecting rod forces will make no significant change to the SCF.The maximum equivalent von Mises stress s3P in the journal fillet is evaluated. The SCF in the journal filletcan be determined in two ways as shown in App.C [3.3.1] and App.C [3.3.2].

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Figure 3 Boundary and load conditions for the pure bending load case

Figure 4 Boundary and load conditions for the 3-point bending load case for an in-line engine

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Figure 5 Load applications for in-line and V-type engines

3.3.1 Method 1This method is analogue to the FVV investigation. The results from 3-point and 4-point bending are combinedas follows:

σ3P = σN3P ∙ βB + σQ3P ∙ βQ

where:

σ3P = as found by the FE calculationσN3P = nominal bending stress in the web centre due to the force F3P [N] applied at the centre-line of the

actual connecting rod, Figure 4 and Figure 5βB = as determined in [3.3]σQ3P = Q3P/(B∙W) where Q3P is the radial (shear) force in the web due to the force F3P applied at the

centre-line of the actual connecting rod, see Sec.1 Figure 3.

3.3.2 Method 2This method is not analogous to the FVV investigation. In a statically determined system as one crank throwsupported on two bearings, the bending moment and radial (shear) force are proportional. Therefore thejournal fillet SCF can be found directly by the 3 point bending FE calculation.The SCF is then calculated according to:

For symbols see method 1.When using this method the radial force and stress determination in M53 becomes superfluous. Thealternating bending stress in the journal fillet as per UR M53 2.1.3 is then evaluated

σBG = ±|βBQ ∙ σBFN|where σBFN is calculated in M53.

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Note that the use of this method does not apply to the crankpin fillet and that this SCF shall not be used inconnection with calculation methods other than those assuming a statically determined system as in M53.

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APPENDIX D GUIDANCE FOR EVALUATION OF FATIGUE TESTS

1 IntroductionFatigue testing can be divided into two main groups; testing of small specimens and full size crank throws.For crankshafts without any fillet surface treatment, the fatigue strength can be determined by testing smallspecimens taken from a full size crank throw. One advantage is the rather high number of specimens whichcan be then manufactured. Another advantage is that the tests can be made with different stress ratios (R-ratios) and/or different modes e.g. axial, bending and torsion, with or without a notch. This is required forevaluation of the material data to be used with critical plane criteria.For crankshafts with surface treatment the fatigue strength can only be determined through testing of fullsize crank throws. For cost reasons this usually means a low number of crank throws. The load can beapplied by hydraulic actuators in a 3- or 4-point bending arrangement, or by an exciter in a resonance testrig. The latter is frequently used, although it usually limits the stress ratio to R = -1.Testing can be made using the staircase method or a modified version thereof which is presented in thisdocument. Other statistical evaluation methods may also be applied.

2 Evaluation of test results

2.1 PrinciplesPrior to fatigue testing, the crankshaft shall be tested as required by quality control procedures, e.g. forchemical composition, mechanical properties, surface hardness, hardness depth and extension, fillet surfacefinish, etc.The test samples should be prepared so as to represent the "lower end" of the acceptance range e.g. forinduction hardened crankshafts this means the lower range of acceptable hardness depth, the shortestextension through a fillet, etc. Otherwise the mean value test results should be corrected with a confidenceinterval: a 90% confidence interval may be used both for the sample mean and the standard deviation.The test results, when applied in M53, shall be evaluated to represent the mean fatigue strength, with orwithout taking into consideration the 90% confidence interval as mentioned above. The standard deviationshould be considered by taking the 90% confidence into account. Subsequently the result to be used as thefatigue strength is then the mean fatigue strength minus one standard deviation.If the evaluation aims to find a relationship between (static) mechanical properties and the fatigue strength,the relation shall be based on the real (measured) mechanical properties, not on the specified minimumproperties.The calculation technique presented in [2.4] was developed for the original staircase method. However, sincethere is no similar method dedicated to the modified staircase method the same is applied for both.

2.2 Staircase methodIn the original staircase method, the first specimen is subjected to a stress corresponding to the expectedaverage fatigue strength. If the specimen survives 107 cycles, it is discarded and the next specimen issubjected to a stress that is one increment above the previous, i.e. a survivor is always followed by the nextusing a stress one increment above the previous. The increment should be selected to correspond to theexpected level of the standard deviation.When a specimen fails prior to reaching 107 cycles, the obtained number of cycles is noted and the nextspecimen is subjected to a stress that is one increment below the previous. With this approach the sum offailures and run-outs is equal to the number of specimens.This original staircase method is only suitable when a high number of specimens are available. Throughsimulations it has been found that the use of about 25 specimens in a staircase test leads to a sufficientaccuracy in the result.

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2.3 Modified staircase methodWhen a limited number of specimens are available, it is advisable to apply the modified staircase method.Here the first specimen is subjected to a stress level that is most likely well below the average fatiguestrength. When this specimen has survived 107 cycles, this same specimen is subjected to a stress levelone increment above the previous. The increment should be selected to correspond to the expected level ofthe standard deviation. This is continued with the same specimen until failure. Then the number of cycles isrecorded and the next specimen is subjected to a stress that is at least 2 increments below the level wherethe previous specimen failed.With this approach the number of failures usually equals the number of specimens. The number of run-outs,counted as the highest level where 107 cycles were reached, also equals the number of specimens.The acquired result of a modified staircase method should be used with care, since some results availableindicate that testing a runout on a higher test level, especially at high mean stresses, tends to increase thefatigue limit. However, this "training effect" is less pronounced for high strength steels (e.g. UTS > 800 MPa).If the confidence calculation is desired or necessary, the minimum number of test specimens is 3.

2.4 Calculation of sample mean and standard deviationA hypothetical example of tests for 5 crank throws is presented further in the subsequent text. When usingthe modified staircase method and the evaluation method of Dixon and Mood, the number of samples will be10, meaning 5 run-outs and 5 failures, i.e.:

Number of samples: n = 10

Furthermore the method distinguishes between

Less frequent event is failure: C = 1

Less frequent event is run-outs: C = 2

The method uses only the less frequent occurrence in the test results, i.e. if there are more failures than run-outs, then the number of run-outs is used, and vice versa.

In the modified staircase method, the number of run-outs and failures are usually equal. However, thetesting can be unsuccessful, e.g. the number of run-outs can be less than the number of failures if aspecimen with 2 increments below the previous failure level goes directly to failure. On the other hand, if thisunexpected premature failure occurs after a rather high number of cycles, it is possible to define the levelbelow this as a run-out.

Dixon and Mood's approach, derived from the maximum likelihood theory, which also may be applied here,especially on tests with few samples, presented some simple approximate equations for calculating thesample mean and the standard deviation from the outcome of the staircase test. The sample mean can becalculated as follows:

when C = 1

when C = 2

The standard deviation can be found by

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Where

Sa0 = the lowest stress level for the less frequent occurrenced = the stress incrementF = ∑fiA = ∑i · fiB = ∑i

2 · fii = the stress level numberingfi = is the number of samples at stress level i

The formula for the standard deviation is an approximation and can be used when

and 0.5 · s < d < 1.5 · s

If any of these two conditions are not fulfilled, a new staircase test should be considered or the standarddeviation should be taken quite large in order to be on the safe side.

If increment d is greatly higher than the standard deviation s, the procedure leads to a lower standarddeviation and a slightly higher sample mean, both compared to values calculated when the differencebetween the increment and the standard deviation is relatively small. Respectively, if increment d is muchless than the standard deviation s, the procedure leads to a higher standard deviation and a slightly lowersample mean.

Guidance note:Example 1 Hypothetical test results may look as shown in Figure 1. The processing of the results and the evaluation of the samplemean and the standard deviation are shown in Figure 2.

Figure 1 Log sheet of a modified staircase test

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Figure 2 Processing of the staircase test results

---e-n-d---of---g-u-i-d-a-n-c-e---n-o-t-e---

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2.5 Confidence interval for mean fatigue limitIf the staircase fatigue test is repeated, the sample mean and the standard deviation will most likely bedifferent from the previous test. Therefore, it is necessary to assure with a given confidence that therepeated test values will be above the chosen fatigue limit by using a confidence interval for the samplemean. The confidence interval for the sample mean value with unknown variance is known to be distributedaccording to the t-distribution (also called student's t-distribution) which is a distribution symmetric aroundthe average.

Figure 3 Student's t-distribution

If Sa is the empirical mean and s is the empirical standard deviation over a series of n samples, in which thevariable values are normally distributed with an unknown sample mean and unknown variance, the (1 - a) ·100% confidence interval for the mean is:

The resulting confidence interval is symmetric around the empirical mean of the sample values, and thelower endpoint can be found as;

which is the mean fatigue limit (population value) to be used to obtain the reduced fatigue limit where thelimits for the probability of failure are taken into consideration.

Guidance note:

Example 2 Applying a 90% confidence interval (α = 0.1) and n = 10 (5 failures and 5 run-outs) leads to tα, n-1 = 1.383, taken froma table for statistical evaluations (E. Doughtery: Probability and Statistics for the Engineering, Computing and Physical Sciences,1990. Note that v = 1 in the tables).Hence:

To be conservative, some authors would consider n to be 5, as the physical number of used specimen, then tα, n-1=1.533.


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2.6 Confidence interval for standard deviationThe confidence interval for the variance of a normal random variable is known to possess a chi-squaredistribution with n – 1 degrees of freedom.

Figure 4 Chi-square distribution

An assumed fatigue test value from n samples has a normal random variable with a variance of σ2 and hasan empirical variance s2. Then a (1 - a) · 100% confidence interval for the variance is:

A (1 - s) · 100% confidence interval for the standard deviation is obtained by the square root of the upperlimit of the confidence interval for the variance and can be found by

The standard deviation (population value) shall be used to obtain the fatigue limit, where the limits for theprobability of failure are taken into consideration.

Guidance note:

Example 3 Applying a 90% confidence interval (α = 0.1) and n = 10 (5 failures and 5 run-outs) leads to = 4.168, taken from a tablefor statistical evaluations (E. Doughtery: Probability and Statistics for the Engineering, Computing and Physical Sciences, 1990).Hence,

To be conservative, some authors would consider n to be 5, as the physical number of the used specimen, then

= 1.064.


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3 Small specimen testingIn this connection a small specimen is considered to be one of the specimens taken from a crank throw.Since the specimens shall be representative for the fillet fatigue strength, they should be taken out close tothe fillets, as shown in Figure 5.

Figure 5 Specimen locations in a crank throw

The (static) mechanical properties shall be determined as stipulated by the quality control procedures.

3.1 Determination of bending fatigue strengthIt is advisable to use un-notched specimens in order to avoid uncertainties related to the stress gradientinfluence. Push-pull testing (stress ratio R = -1) is preferred, but for the purpose of critical plane criteriaanother stress ratio may be added.A. If the objective of the testing is to document the influence of high cleanliness, test samples taken frompositions approximately 120 degrees in a circumferential direction may be used. See Sec.3 Figure 1.B. If the objective of the testing is to document the influence of continuous grain flow (cgf) forging, thespecimens should be restricted to the vicinity of the crank plane.

3.2 Determination of torsional fatigue strengthA. If the specimens are subjected to torsional testing, the selection of samples should follow the sameguidelines as for bending above. The stress gradient influence shall be considered in the evaluation.B. If the specimens are tested in push-pull, the samples should be taken out at an angle of 45 degrees to thecrank plane. When taking the specimen at a distance from the (crank) middle plane of the crankshaft alongthe fillet, this plane rotates around the pin center point making it possible to resample the fracture directiondue to torsion (the results shall be converted into the pertinent torsional values.)

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3.3 Other test positionsIf the test purpose is to find fatigue properties and the crankshaft is forged in a manner likely to lead tocgf, the specimens may also be taken longitudinally from a prolonged shaft piece where specimens formechanical testing are usually taken. The condition is that this prolonged shaft piece is heat treated as a partof the crankshaft and that the size is so as to result in a similar quenching rate as the crank throw.When using test results from a prolonged shaft piece, it shall be considered how well the grain flow in thatshaft piece is representative for the crank fillets.

3.4 Correlation of test resultsWhen using the bending fatigue properties from tests mentioned in this section, it should be kept in mindthat successful continuous grain flow (cgf) forging leading to elevated values compared to other (non cgf)forging (or empirical approaches), will normally not lead to a torsional fatigue strength improvement ofthe same magnitude. In such cases it is advised to either carry out also torsional testing or to make aconservative assessment of the torsional fatigue strength, e.g. by using no credit for cgf. This approach isapplicable when using the Gough Pollard criterion. However, this approach is not recognised when using thevon Mises or a multi-axial criterion such as Findley.If the found ratio between bending and torsion fatigue differs from V3, one should consider replacing theuse of the von Mises criterion with the Gough Pollard criterion. Also, if critical plane criteria are used, it shallbe kept in mind that cgf makes the material inhomogeneous in terms of fatigue strength, meaning that thematerial parameters differ with the directions of the planes.Any addition of influence factors shall be made with caution. If for example a certain addition for superclean steel is documented, it may not necessarily be fully combined with a K-factor for cgf. Direct testing ofsamples from a super clean and cgf forged crank is preferred.

4 Full size testing

4.1 Hydraulic pulsationA hydraulic test rig can be arranged for testing a crankshaft in 3-point or 4-point bending as well as intorsion. This allows for testing with any R-ratio.Although the applied load should be verified by strain gauge measurements on plain shaft sections for theinitiation of the test, it is not necessarily used during the test for controlling load. It is also pertinent to checkfillet stresses with strain gauge chains. Furthermore, it is important that the test rig provides boundaryconditions as defined in App.C.The (static) mechanical properties shall be determined as stipulated by the quality control procedures.

4.2 Resonance testerA rig for bending fatigue normally works with an R-ratio of -1. Due to operation close to resonance, theenergy consumption is moderate. Moreover, the frequency is usually relatively high, meaning that 107 cyclescan be reached within some days. Figure 6 shows a layout of the testing arrangement. The applied loadshould be verified by strain gauge measurements on plain shaft sections. It is also pertinent to check filletstresses with strain gauge chains.Clamping around the journals shall be arranged in a way that prevents severe fretting which could leadto a failure under the edges of the clamps. If some distance between the clamps and the journal fillets isprovided, the loading is consistent with 4-point bending and thus representative for the journal fillets also.In an engine the crankpin fillets normally operate with an R-ratio slightly above -1 and the journal filletsslightly below -1. If found necessary, it is possible to introduce a mean load (deviate from R = -1) by meansof a spring pre-load.

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Figure 6 A testing arrangement of the resonance tester for bending loading

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Figure 7 A testing arrangement of the resonance tester for torsion loading

A rig for torsion fatigue can also be arranged as shown in Figure 7. When a crank throw is subjected totorsion, the twist of the crankpin makes the journals move sideways. If one single crank throw is tested ina torsion resonance test rig, the journals with their clamped-on weights will vibrate heavily sideways. Thissideways movement of the clamped-on weights can be reduced by having two crank throws, especially if thecranks are almost in the same direction. However, the journal in the middle will move more.Since sideway movements can cause some bending stresses, the plain portions of the crankpins shouldalso be provided with strain gauges arranged to measure any possible bending that could have aninfluence on the test results. Similarly to the bending case the applied load shall be verified by strain gaugemeasurements on plain shaft sections. It is also pertinent to check fillet stresses with strain gauge chains aswell.

5 Use of existing results for similar crankshaftsFor fillets or oil bores without surface treatment, the fatigue properties found by testing may be used forsimilar crankshaft designs providing:

— The same material type— Cleanliness on the same or better level— The same mechanical properties can be granted (size versus hardenability)— Similar manufacturing process

Induction hardened or gas nitrited crankshafts will suffer fatigue either at the surface or at the transition tothe core. The surface fatigue strength as determined by fatigue tests of full size cranks, may be used on an

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equal or similar design as the tested crankshaft when the fatigue initiation occurred at the surface. With thesimilar design it is meant that the same material type and surface hardness are used and the fillet radius andhardening depth are within approximately ± 30% of the tested crankshaft.Fatigue initiation in the transition zone can be either subsurface, i.e. below the hard layer, or at the surfacewhere the hardening ends. The fatigue strength at the transition to the core can be determined by fatiguetests as described above, provided that the fatigue initiation occurred at the transition to the core. Testsmade with the core material only will not be representative since the tension residual stresses at thetransition are lacking.For cold rolled or stroke peened crankshafts, the results obtained by one full-size crank test can be appliedto another crank size, provided that the base material is of the same type and that the treatment is made inorder to obtain the same level of compressive residual stresses at the surface as well as in the depth. Thismeans that both the extension and the depth of rolling or peening shall be proportional to the fillet radius.It shall be noted also what some recent research has shown: The fatigue limit can decrease in the very highcycle domain with subsurface crack initiation due to trapped hydrogen that accumulates through diffusionaround some internal defect functioning as an initiation point. In these cases it would be appropriate toreduce the fatigue limit by some percent per decade of cycles beyond 107. Based on a publication by YukitakaMurakami "Metal Fatigue: Effects of Small Defects and Nonmetallic Inclusions" the reduction is suggested tobe 5% per decade.

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APPENDIX E SURFACE TESTED FILLETS AND OIL BORE OUTLETS

1 IntroductionThis appendix deals with surface treated fillets and oil bore outlets. The various treatments are explainedand some empirical formulae are given for calculation purposes. Conservative empiricism has been appliedintentionally, in order to be on the safe side from a calculation standpoint.Please note that measurements or more specific knowledge should be used if available. However, in the caseof a wide scatter (e.g. for residual stresses) the values should be chosen from the end of the range thatwould be on the safe side for calculation purposes.

2 Definition of surface treatment‘Surface treatment’ is a term covering treatments such as thermal, chemical or mechanical operations,leading to inhomogeneous material properties – such as hardness, chemistry or residual stresses – from thesurface to the core.

2.1 Surface treatment methodsThe following list covers possible treatment methods and how they influence the properties that are decisivefor the fatigue strength.

Table 1 Surface treatment methods and the characteristics they affect

Treatment method Affecting

Induction hardening Hardness and residual stresses

Nitriding Chemistry, hardness and residual stresses

Case hardening Chemistry, hardness and residual stresses

Die quenching (no temper) Hardness and residual stresses

Cold rolling Residual stresses

Stroke peening Residual stresses

Shot peening Residual stresses

Laser peening Residual stresses

Ball coning Residual stresses

It is important to note that since only induction hardening, nitriding, cold rolling and stroke peening areconsidered relevant for marine engines; other methods are not dealt with in this document. In addition diequenching can be considered in the same way as induction hardening.

3 Calculation principlesThe basic principle is that the alternating working stresses shall be below the local fatigue strength (includingthe effect of surface treatment) wherein non-propagating cracks may occur, see also App.E [6.1] for details.This is then divided by a certain safety factor. This applies through the entire fillet or oil bore contour as wellas below the surface to a depth below the treatment-affected zone – i.e. to cover the depth all the way to thecore.

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Consideration of the local fatigue strength shall include the influence of the local hardness, residual stressand mean working stress. The influence of the ‘giga-cycle effect’, especially for initiation of subsurface cracks,should be covered by the choice of safety margin.It is of vital importance that the extension of hardening/peening in an area with concentrated stresses beduly considered. Any transition where the hardening/peening is ended is likely to have considerable tensileresidual stresses. This forms a ‘weak spot’ and is important if it coincides with an area of high stresses.Alternating and mean working stresses shall be known for the entire area of the stress concentration as wellas to a depth of about 1.2 times the depth of the treatment. The following figure indicates this principle inthe case of induction hardening. The base axis is either the depth (perpendicular to the surface) or along thefillet contour.

Figure 1 General principles of stresses as functions of depth

The acceptability criterion shall be applied stepwise from the surface to the core as well as from the point ofmaximum stress concentration along the fillet surface contour to the web.

3.1 Evaluation of local fillet stressesIt is necessary to have knowledge of the stresses along the fillet contour as well as in the subsurface toa depth somewhat beyond the hardened layer. Normally this will be found via FEA as described in App.C.However, the element size in the subsurface range will have to be the same size as at the surface. Forcrankpin hardening only the small element size will have to be continued along the surface to the hard layer.If no FEA is available, a simplified approach may be used. This can be based on the empirically determinedstress concentration factors (SCFs), as in M53.3 if within its validity range, and a relative stress gradientinversely proportional to the fillet radius. Bending and torsional stresses shall be addressed separately. Thecombination of these is addressed by the acceptability criterion.The subsurface transition-zone stresses, with the minimum hardening depth, can be determined by meansof local stress concentration factors along an axis perpendicular to the fillet surface. These functions ĮB-localand ĮT-local have different shapes due to the different stress gradients.

The SCFs αβ and αT are valid at the surface, the local αβ-local and αT-local drop with increasing depth. Therelative stress gradients at the surface depend on the kind of stress raiser, but for crankpin fillets they can besimplified to 2/RH in bending and 1/RH in torsion. The journal fillets are handled analogously by using RG. The

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nominal stresses are assumed to be linear from the surface to a midpoint in the web between the crankpinfillet and the journal fillet for bending and to the crankpin or journal center for torsion.The local SCFs are then functions of depth t according to Equation E-1 as shown in Figure 2 for bending andrespectively for torsion in Equation E-2 and Figure 3.

Equation E-1

Equation E-2

Figure 2 Bending SCF in the crankpin fillet as a function of depth

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Figure 3 Torsional SCF in the crankpin fillet as a function of depth

In Figure 2 the corresponding SCF for the journal fillet can be found by replacing RH with RG. In Figure 3 thecorresponding SCF for the journal fillet can be found by replacing RH with RG and D with DG.If the pin is hardened only and the end of the hardened zone is closer to the fillet than three times themaximum hardness depth, FEA should be used to determine the actual stresses in the transition zone.

3.2 Evaluation of oil bore stressesStresses in the oil bores can be determined also by FEA. The element size should be less than 1/8 of the oilbore diameter DO and the element mesh quality criteria should be followed as prescribed in App.C. The fineelement mesh should continue well beyond a radial depth corresponding to the hardening depth.The loads to be applied in the FEA are the torque, App.C [3.1] – and the bending moment, with four-pointbending as in App.C [3.2].If no FEA is available, a simplified approach may be used. This can be based on the empirically determinedSCF from M53.3 if within its applicability range. Bending and torsional stresses at the point of peak stressesare combined as in M53.5.

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Figure 4 Stresses and hardness in induction hardened oil holes

Figure 4 indicates a local drop of the hardness in the transition zone between a hard and soft material.Whether this drop occurs depends also on the tempering temperature after quenching in the QT process.The peak stress in the bore occurs at the end of the edge rounding. Within this zone the stress dropsalmost linearly to the center of the pin. As can be seen from Figure 4, for shallow (A) and intermediate (B)hardening, the transition point practically coincides with the point of maximal stresses. For deep hardeningthe transition point comes outside of the point of peak stress and the local stress can be assessed as aportion (1-2·tH/D) of the peak stresses where tH is the hardening depth.The subsurface transition-zone stresses (using the minimum hardening depth) can be determined by meansof local stress concentration factors along an axis perpendicular to the oil bore surface. These functions γB-

local and γT-local have different shapes, because of the different stress gradients.

The stress concentration factors γB and γT are valid at the surface, the local SCFs γB-local and γT-local dropwith increasing depth. The relative stress gradients at the surface depend on the kind of stress raiser, butfor crankpin oil bores they can be simplified to 4/DO in bending and 2/DO in torsion. The local SCFs are thenfunctions of the depth t:

Equation E-3

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Equation E-4

3.3 Acceptability criteriaAcceptance of crankshafts is based on fatigue considerations; M53 compares the equivalent alternating stressand the fatigue strength ratio to an acceptability factor of Q ≥ 1.15 for oil bore outlets, crankpin fillets andjournal fillets. The term ‘safety factor’ is intentionally not used, since the method contains several so-calledhidden safeties.Some assumptions leading to so-called hidden safety:

i) Analytical equations do not cover complicated crank web geometries or fillet shapesii) Three-point bending applies on one crank throw without influence from adjacent cranksiii) Full torque applies on the crankpin, i.e. no restraint in main bearings is includediv) Stress concentration for torsion and bending occurs at the same location in a filletv) Maximum torsion and maximum bending stresses are in phasevi) Fatigue strength is assessed without the influence of mean stress

For calculations assuming i) to iii), even if SCFs are found by FEA, the acceptability factor of 1.15 shouldremain. When more accurate calculations are made (e.g. when one or more of the hidden safeties arebypassed) the acceptability factor turns more toward the safety factor and a suitable value and approachshould be applied.

4 Induction hardeningGenerally the hardness specification shall specify the surface hardness range i.e. minimum and maximumvalues, the minimum and maximum extension in or through the fillet and also the minimum and maximumdepth along the fillet contour.The induction hardening depth is defined as the depth where the hardness is 80% of the minimum specifiedsurface hardness.

Figure 5 Typical hardness as a function of depth

In Figure 5 the arrows indicate the defined hardening depth. Note the indicated potential hardness drop atthe transition to the core. This can be a weak point as local strength may be reduced and tensile residual

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stresses may occur. In the case of crankpin or journal hardening only, the minimum distance to the fillet shallbe specified due to the tensile stress at the heat-affected zone as shown in Figure 6.

Figure 6 Residual stresses along the surface of a pin and fillet

If the hardness-versus-depth profile and residual stresses are not known or specified, one may assume thefollowing:

— the hardness profile consists of two layers (Figure 5):

— constant hardness from the surface to the transition zone— constant hardness from the transition zone to the core material

— residual stresses in the hard zone of 200 MPa (compression)

— transition-zone hardness as 90% of the core hardness unless the local hardness drop is avoided

— transition-zone maximum residual stresses (von Mises) of 300 MPa tension.

If the crankpin or journal hardening ends close to the fillet, the influence of tensile residual stresses shall beconsidered. If the minimum distance between the end of the hardening and the beginning of the fillet is morethan 3 times the maximum hardening depth, the influence may be disregarded.

4.1 Local fatigue strengthInduction-hardened crankshafts will suffer fatigue either at the surface or at the transition to the core. Thefatigue strengths, for both the surface and the transition zone, can be determined by fatigue testing of fullsize cranks as described in App.D. In the case of a transition zone, the initiation of the fatigue can be eithersubsurface (i.e. below the hard layer) or at the surface where the hardening ends. Tests made with thecore material only will not be representative since the tensile residual stresses at the transition are lacking.Alternatively, the surface fatigue strength (principal stress) can be determined empirically as follows whereHV is the surface (Vickers) hardness. Equation E-5 provides a conservative value, with which the fatiguestrength is assumed to include the influence of the residual stress. The resulting value is valid for a workingstress ratio of R = -1:

[MPa] Equation E-5

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It shall be noted also that the mean stress influence of induction-hardened steels may be significantly higherthan that for QT steels.The fatigue strength in the transition zone, without taking into account any possible local hardness drop, shallbe determined by the equation introduced in UR M53.6. For journal and respectively to crankpin fillet applies:

Equation E-6

Where

Y = DG and X = RG for journal fillet

Y = D and X = RH for crankpin fillet

The influence of the residual stress is not included in Equation E-6.For the purpose of considering subsurface fatigue, below the hard layer, the disadvantage of tensile residualstresses shall be considered by subtracting 20% from the value determined above. This 20% is based on themean stress influence of alloyed quenched and tempered steel having a residual tensile stress of 300 MPa.When the residual stresses are known to be lower, also smaller value of subtraction shall be used. For low-strength steels the percentage chosen should be higher.For the purpose of considering surface fatigue near the end of the hardened zone – i.e. in the heat-affectedzone shown in the Figure 6 – the influence of the tensile residual stresses can be considered by subtracting acertain percentage, in accordance with Table 2, from the value determined by the above formula.

Table 2 The influence of tensile residual stresses at a given distance from the end of thehardening

5 NitridingThe hardness specification shall include the surface hardness range (min and max) and the minimum andmaximum depth. Only gas nitriding is considered.

The depth of the hardening is defined in different ways in the various standards and the literature. The mostpractical method to use in this context is to define the nitriding depth tN as the depth to a hardness of 50 HVabove the core hardness.

The hardening profile should be specified all the way to the core. If this is not known, it may be determinedempirically via the following formula:

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Equation E-7

Where

t = the local depthHV(t) = hardness at depth tHVcore = core hardnessHVsurface = surface hardnessTN = nitriding depth as defined above

5.1 Local fatigue strengthIt is important to note that in nitrided crankshaft cases, fatigue is found either at the surface or at thetransition to the core. This means that the fatigue strength can be determined by tests as described in App.DAlternatively, the surface fatigue strength (principal stress) can be determined empirically and conservativelyas follows. This is valid for a surface hardness of 600HV or greater:

σFsurface = 450 MPaNote that this fatigue strength is assumed to include the influence of the surface residual stress and appliesfor a working stress ratio of R = -1.The fatigue strength in the transition zone can be determined by the equation introduced in UR M53.6. Forcrankpin and respectively to journal applies:

Equation E-8

Where

Y = DG and X =RG

for journal fillet

Y = D and X =RH

for crankpin fillet

The influence of the residual stress is not included in Equation E-8.In contrast to induction-hardening the nitrited components have no such distinct transition to the core.Although the compressive residual stresses at the surface are high, the balancing tensile stresses in thecore are moderate because of the shallow depth. For the purpose of analysis of subsurface fatigue thedisadvantage of tensile residual stresses in and below the transition zone may be even disregarded in view ofthis smooth contour of a nitriding hardness profile.Nevertheless, in principle the calculation should be carried out along the entire hardness profile and it canbe limited to a simplified approach of examining the surface and an artificial transition point. This artificialtransition point can be taken at the depth where the local hardness is approximately 20 HV above the corehardness. In such a case, the properties of the core material should be used. This means that the stresses at

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the transition to the core can be found by using the local SCF formulae mentioned earlier when inserting t =1.2 · tN.

Figure 7 Sketch of the location for the artificial transition point in the depth direction

6 Cold formingThe advantage of stroke peening or cold rolling of fillets is the compressive residual stresses introduced inthe high-loaded area. Even though surface residual stresses can be determined by X-ray diffraction techniqueand subsurface residual stresses can be determined through neutron diffraction, the local fatigue strength isvirtually non-assessable on that basis since suitable and reliable correlation formulae are hardly known.Therefore the fatigue strength shall be determined by fatigue testing; see also App.D. Such testing isnormally carried out as four-point bending, with a working stress ratio of R = –1. From these results thebending fatigue strength, surface- or subsurface-initiated depending on the manner of failure, can bedetermined and expressed as the representative fatigue strength for applied bending in the fillet.In comparison to bending, the torsion fatigue strength in the fillet may differ considerably from the ratio(utilized by the von Mises criterion). The forming affected depth that is sufficient to prevent subsurfacefatigue in bending, may still allow subsurface fatigue in torsion. Another possible reason for the difference inbending and torsion could be the extension of the highly stressed area.The results obtained in a full size crank test can be applied for another crank size provided that the basematerial (alloyed Q + T) is of the same type and that the forming is done so as to obtain the same levelof compressive residual stresses at the surface as well as through the depth. This means that both theextension and the depth of the cold forming shall be proportional to the fillet radius.

6.1 Stroke peening by means of a ballThe fatigue strength obtained can be documented by means of full size crank tests or by empirical methods ifapplied on the safe side. If both bending and torsion fatigue strengths have been investigated and differ fromthe ratio, the von Mises criterion should be excluded.If only bending fatigue strength has been investigated, the torsional fatigue strength should be assessedconservatively. If the bending fatigue strength is concluded to be x% above the fatigue strength of the non-peened material, the torsional fatigue strength should not be assumed to be more than 2/3 of x% above thatof the non-peened material.As a result of the stroke peening process the maximum of the compressive residual stress is found in thesubsurface area. Therefore, depending on the fatigue testing load and the stress gradient, it is possibleto have higher working stresses at the surface in comparison to the local fatigue strength of the surface.Because of this phenomenon small cracks may appear during the fatigue testing, which will not be ableto propagate in further load cycles and/or with further slight increases of the testing load because of theprofile of the compressive residual stress. Put simply, the high compressive residual stresses below the

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surface ‘arrest’ small surface cracks. This is illustrated in Figure 8 as gradient load 2. Straight lines 1, 2 and 3represent different possible load stress gradients.

Figure 8 Working and residual stresses below the stroke-peened surface

In fatigue testing with full-size crankshafts these small “hairline cracks” should not be considered to be thefailure crack. The crack that is technically the fatigue crack leading to failure, and that therefore shuts offthe test-bench, should be considered for determination of the failure load level. This also applies if induction-hardened fillets are stroke-peened.In order to improve the fatigue strength of induction-hardened fillets it is possible to apply the strokepeening process in the crankshafts’ fillets after they have been induction-hardened and tempered to therequired surface hardness. If this is done, it might be necessary to adapt the stroke peening force to thehardness of the surface layer and not to the tensile strength of the base material. The effect on the fatiguestrength of induction hardening and stroke peening the fillets shall be determined by a full size crankshafttest. See App.D.Fatigue results from tests on one crankshaft may be used for a similar crankshaft if all of the followingcriteria are fulfilled:

— ball size relative to fillet radius within 90 ± 5%— at least the same circumferential extension of the stroke peening— at least the same angular extension of fillet contour stroke-peened— similar base material, e.g. alloyed quenched and tempered— forward feed of ball of the same proportion of the radius— force applied to ball proportional to base material hardness (if different)— force applied to ball proportional to square of ball radius.

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6.2 Cold rollingThe fatigue strength can be obtained by means of full size crank tests or by empirical methods if these areapplied so as to be on the safe side. If both bending and torsion fatigue strengths have been investigated,and differ from the ratio, the von Mises criterion should be excluded.If only bending fatigue strength has been investigated, the torsional fatigue strength should be assessedconservatively. If the bending fatigue strength is concluded to be x% above the fatigue strength of the non-rolled material, the torsional fatigue strength should not be assumed to be more than 2/3 of x% above thatof the non-rolled material.Fatigue test results from testing of a crankshaft may be applied to another crankshaft if that crankshaft issimilar and all of the following criteria are fulfilled:

— at least the same circumferential extension of cold rolling— at least the same angular extension of fillet contour rolled— similar base material, e.g. alloyed quenched and tempered— roller force to be calculated so as to achieve at least the same relative (to fillet radius) depth of treatment

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APPENDIX F STRESS CONCENTRATION FACTORS IN THE OILBOREOUTLETS OF CRANKSHAFTS BY UTILIZING FINITE ELEMENTMETHOD

1 GeneralThe objective of the analysis is to substitute the analytically calculated stress concentration factor (SCF)at the oilbore outlet by suitable finite element method (FEM) calculated figures. The analytical method isbased on empirical formulae developed from strain gauge or photo-elasticity measurements of various roundbars. Use of these formulae beyond any of the various validity ranges can lead to erroneous results in eitherdirection, i.e. results that more inaccurate than indicated by the mentioned standard deviations. Thereforethe FEM-based method is highly recommended.The SCF calculated according to the rules of this document are defined as the ratio of stresses calculated byFEM to nominal stresses calculated analytically. When used in connection with the present method in M53 vonMises stresses shall be calculated for bending and principal stresses for torsion.The analysis shall be conducted as linear elastic FE analysis, and unit loads of appropriate magnitude shall beapplied for all load cases.It is advised to check element accuracy of the FE solver in use, e.g. by modelling a simple geometry andcomparing the stresses obtained by FEM with the analytical solution.Boundary element method (BEM) may be used instead of FEM.

2 Model requirementsThe basic recommendations and perceptions for building the FE-model are presented in [2.1]. It is obligatoryfor the final FE-model to fulfil a requirement in [2.3].

2.1 Element mesh recommendationsIn order to fulfil the mesh quality criteria it is advised to construct the FE model for the evaluation of StressConcentration Factors according to the following recommendations:

— the model consists of one complete crank, from the main bearing centerline to the opposite side mainbearing centerline

— element types used in the vicinity of the outlets:

— 10 node tetrahedral elements— 8 node hexahedral elements— 20 node hexahedral elements

— mesh properties in oilbore outlet:

— maximum element size a = r / 4 through the entire outlet fillet as well as in the bore direction.(If 8 node hexahedral elements are used even smaller element size is required to meet the qualitycriterion.)

— recommended manner for element size in fillet depth direction

— first layer thickness equal to element size of a— second layer thickness equal element to size of 2a— third layer thickness equal element to size of 3a

— generally the rest of the crank should be suitable for numeric stability of the solver

— drillings and holes for weight reduction shall be modelled.

Sub-modeling may be used as far as the software requirements are fulfilled.

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2.2 MaterialUR M53 does not consider material properties such as Young’s Modulus (E) and Poisson’s ratio (n). In FEanalysis those material parameters are required, as primarily strain is calculated and stress is derived fromstrain using the Young’s Modulus and Poisson’s ratio. Reliable values for material parameters shall be used,either as quoted in literature or measured on representative material samples.

For steel the following is advised: E= 2.05·105 MPa and ν = 0.3.

2.3 Element mesh quality criteriaIf the actual element mesh does not fulfil any of the following criteria at the examined area for SCFevaluation, then a second calculation with a refined mesh shall be performed.

2.3.1 Principal stresses criterionThe quality of the mesh should be assured by checking the stress component normal to the surface of theoil bore outlet radius. Ideally, this stress should be zero. With principal stresses σ1, σ2 and σ3 the followingcriterion is required:

min(|σ1|,|σ2|,|σ3|) < 0.03 · max(|σ1|,|σ2|,|σ3|)

2.3.2 Averaged/un-averaged stresses criterionThe criterion is based on observing the discontinuity of stress results over elements at the fillet for thecalculation of SCF:

— Unaveraged nodal stress results calculated from each element connected to a nodei should differ less thanby 5% from the 100% averaged nodal stress results at this nodei at the examined location.

3 Load casesTo substitute the analytically determined SCF in UR M53 the following load cases shall be calculated.

3.1 TorsionThe structure is loaded in pure torsion. In the model surface warp at the end faces is suppressed. Torque isapplied to the central node located at the crankshaft axis. This node acts as the master node with 6 degreesof freedom and is connected rigidly to all nodes of the end face.Boundary and load conditions are valid for both in-line- and V- type engines.

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Figure 1 Boundary and load conditions for the torsion load case

For all nodes in oilbore outlet, principal stresses are extracted and the equivalent torsional stress iscalculated:

The maximum value is taken for the subsequent calculation of the SCF:

Where nominal stress referred to the crankpin as per UR M53 2.2.2:

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3.2 BendingThe structure is loaded in pure bending. In the model surface warp at the end faces is suppressed.The bending moment is applied to the central node located at the crankshaft axis. This node acts as masternode with 6 degrees of freedom and is connected rigidly to all nodes of the end face.Boundary and load conditions are valid for both in-line- and V- type engines.

Figure 2 Boundary - and load conditions for the pure bending load case

For all nodes in oil bore outlet von Mises equivalent stresses σequiv are extracted. The maximum value is usedto calculate the SCF according to:

The nominal stress sN is calculated as per UR M53 2.1.2.2 with the bending moment M:

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Documents

DNVGL-CG-0037 Calculation of crankshafts for reciprocating ......Changes - current Class guideline — DNVGL-CG-0037. Edition November 2015 Page 3 Calculation of crankshafts for reciprocating