56 Professional BoatBuilder

highlight some of the common causes of failure in aluminum boats that are too often overlooked during planning and construction.

To better understand why such mis-takes are made and how to correct them, let’s take a look at a common fatigue failure in a series of high-speed aluminum powerboats I worked on in the early 1990s: cracking in the alumi-num structure in transoms and around waterjets.

At the time, the popularity of water-jet propulsion was booming, and with it came a troubling rise in these cracks. They were common enough to almost be expected. The problems have now been well documented and,

Creating the complex structure and graceful curves of an aluminum

boat requires the profound transfor-mation of fl at, square, virgin plate. It is rolled, bent, forged, bashed, scored, heated, welded into shape, and painted. Done incorrectly, any one of these processes can seriously compro-mise the fatigue life of the resulting structure. Such failings can be exacer-bated by original designs that over-look the simple details needed to accommodate the properties and peculiarities of the material.

The relationship between aluminum fabrication practices and fatigue failure is complex. While I cannot possibly cover all its aspects in a brief article, I’ll

Above—Good welding practices coupled with sound design of structural elements go a long way in minimizing the fatigue failure that’s common in many aluminum boats, especially those that see hard service in rough conditions.

Fabrication and Fatigue Failure in AluminumClose attention to design and best practices can simplify construction and extend the working life of a high-speed aluminum boat.

by John KecsmarGraphics courtesy of the author

(except where noted)

June/July 2012 57

In contrast, aluminum has no fatigue limit. Constant exposure of aluminum to repeated loading will ultimately lead to failure, no matter how low the actual stress magni-tude experienced for all practical purposes.

Faced with numerous failures, we were beginning to think that alumi-num was a poor choice for a light-weight boatbuilding material; it seemed cursed. It turns out that many of our problems—and those of other builders—were caused by applying steel design and building methods to aluminum, a very different metal. Steel, a forgiving material, has been around for centuries, and the techniques and procedures for work-ing it are highly developed and widely employed. As a boatbuilding material, aluminum is just decades old, so correct, consistent aluminum-fabrication methods are not as com-monly shared. It is imperative that we teach best practices for aluminum to fabricators, especially those who are accustomed to building in steel.

Basic Aluminum Fabrication During fabrication, steel is not

affected and/or changed to the same degree as aluminum, but aluminum has the undeniable advantages of light weight and resistance to cor-rosion. So, how do we reconcile these differences? It’s important for designers and fabricators to recog-nize aluminum’s characteristics so they can maximize its desirable properties, and avoid the sort of failures that plagued our waterjet structures and transoms.

Inspection of the damaged boats revealed classic signs of fatigue failure around the load paths identifi ed in the FEA.

Fatigue In this context, fatigue is best

described by metal fatigue expert L.P. Pook: “a failure of a metal under repeated or otherwise varying load which never reaches a level suffi cient to cause failure in a single applica-tion.” In simple terms: When you apply a force to a structural member, it will bend and become stressed. So long as the stress from static loading is below the material’s yield stress, the member will not fail. If you now repeat this loading many times, although each specifi c load applied does not overstress the member, the simple repetition can initiate a failure owing to the dynamic or cyclic nature of the loading.

In troubleshooting our transom failures, or any fatigue failure of an aluminum boat, there’s an important dis-tinction to be made between fatigue in steel and fatigue in alumi-num. Below a particu-lar level of stress, steel reaches i t s fa t igue limit. No damage or loss of strength will o c cu r be l ow t h a t fatigue limit, regard-less of the number of cycles (Fig. 2).

in most cases, remedied. Builders and operators discovered that corrective action early on could prevent the need for more costly repairs, which in turn prompted more research into these fatigue-related failures.

For the teams designing, building, and repairing these boats, the imme-diate questions were: Why were the boats cracking? And while it was clear that the structural cracks were related to vibration of the waterjets, engine girders were subjected to similar vibra-tions: Why weren’t they failing, too?

Answers weren’t self-evident, so before specifying effective repairs that wouldn’t repeat the problems in the original structure, engineers were tasked with performing a detailed analysis of the mechanics of the failures by applying very basic fi nite element analysis (FEA). Then they strain-gauged a waterjet transom and structure during sea trails to validate the FEA model.

Today, understanding a structure subjected to a range of stresses is readily possible through even more sophisticated FEA, which is often standard at the design phase (see “FEA,” Professional BoatBuilder No. 78, and “Fine-tuning with FEA,” PBB No. 133). Figure 1 shows a typical FEA model of a transom and waterjet structure. This analysis indicated loca-tions of higher stress, though not nec-essarily unacceptably high. Areas identifi ed on the FEA were consistent with locations of the actual failures, but the stresses were all below the designed allowable yield stress of the structure. So why did they fail?

Figure 3-1. Fatigue Limits of Steel and Aluminum

Sem

i-ran

ge o

f Str

ess

x 10 M

Pa

Endurance (Cycles)

30

27

24

21

18

15105 106 107 108 109

Mild steel normalized

Al Cu alloy 2014A T6

Endurance strengthat 50 million cycles

Figure 2. Fatigue Limits of Steel and Aluminum

Figure 1. FEA Model of Transom and Waterjet

58 Professional BoatBuilder

Material-specific quality assurance (QA) is the key to consistently successful design and construction of

aluminum boats. Whether you’re the designer, the plater, or the welder, your awareness of the alloy’s physical properties and limitations is the first step in recognizing what can compromise an aluminum structure, what can go wrong in build, and ultimately how you can avoid such faults. That understanding also clarifies the necessity of different fabrication processes for steel and

aluminum, and helps you avoid the common errors caused by not differentiating between the two materials.

You may ask, “Is strict quality control really that neces-sary? I’m only building a small yacht; won’t it add time and cost? Surely this is only for large commercial builds.” Regardless of size or complexity, once the boat enters service, simple unforeseen vibrations from numerous sources will expose any flaws unintentionally built into the hull due to lack of attention to detail during design

Quality Control

Figure 5. Effects of Notching or Scoring

Max

imum

Str

ess

(ksi

)

45

40

35

30

25

20

15

10

100

0

200

300

103 104 105 106 107 108 109

5

0

Life (Cycles)

MP

a

Lab air

Salt water

Kt = 3

Kt = 12

Lab air—smooth (as-received)Lab air—Kt = 3 edge notchedLab air—Kt = 12 edge notchedSalt water—smooth6061-T4 sheet specimens, R = 0.1

Figure 4. Rotary-Saw Cutting Temperatures in Chip and Plate

Temperatures°F (°C)

tc = 0.036" (0.91mm)

t = 0.024" (0.61mm)

Z

Y

X

1,148 (620)

1,166 (630)

1,184 (640)

1,202 (650)

1,220 (660)

1,274(690)

1,238(670)

1,256(680)

1,292 (700)1,310 (710)

1,328 (720)1,328

1,346 (730)

1,364 (740)

1,382 (750)

Heat

Figure 3. Rotary-Saw Cutting Temperatures in Chip and Plate

Figure 4. Effects of Notching or Scoring

Let’s start with a few fabrication practices that best accommodate aluminum’s principal characteristics.

Avoid marking the plate with scrib-ing tools; these leave marks on the surfaces and can create a slight flaw where a crack can develop and become the site of crack initiation under high loads. Similarly, don’t write on aluminum stock with pencil, as the carbon in it is higher up (more noble) on the galvanic corrosion series.

If aluminum is too coarsely cut, its rough surface can result in potential sites of crack initiation. Cutting with a rotary saw can create considerable heat at the blade tips (Fig. 3). This heat buildup can locally reduce the strength of heat-treatable 6000-series aluminum as well as annealing a strain-hardened 5000-series non-heat-treatable alloy.

Another desirable property of alu-minum that differentiates it from steel is that it requires no special tools for bending. It is important to ensure that the bending tools are free of irregular-ities, which could cause marks or score the surface of the plate. The effects that notching or scoring have on aluminum are shown in Fig. 4. Ranging from minor to major, notches in the plate can reduce the fatigue strength by as much as 75% from its pristine condition.

Clearly, the way you handle, machine, form, bend, and strike the alloy will affect its fatigue strength. Small steel particles from fabrication tools—such as a steel hammer or steel rollers—could unintentionally be embedded into the plate, where they can act as a site for crack propagation or suffer from corrosion leading to stress-corrosion cracking. Always ensure clean, smooth working sur-faces free from contaminants.

June/July 2012 59

this reason the United Kingdom Ministry of Defense accepts only O- temper alloys, which are the softest and most ductile. A typical value of proof stress of 5083 alloy in the O temper is nominally 120–140 MPa (Fig. 5). Strain-hardening the same 5083 alloy to a temper of H2 increases this proof stress to roughly 240 MPa. That sounds good, as the allowable design stress limit has increased. But closer inspection reveals that the same strain-hardening has reduced the margin from yield to ultimate tensile strength (UTS), or failure.

Fig. 6 i l lustrates that strain-hardening each temper also increases the UTS, but at what cost? Thus, the UTS of O temper is approximately 300–330 MPa, whereas in H9 temper, the UTS is 420. As strain hardening increases, the percentage over the base yield proof stress the temper can carry, decreases. It ranges from 107% for O temper, down to a paltry 14% in the H9 temper.

This means that if the alloy is over-loaded from its design-allowable stress in O temper, the amount of reserve capacity is 107%, or double the design limit—an inbuilt safety fac-tor of 2. But in the H9 temper, for example, the margin from yield to overload is so small that an overload can cause instant failure of the struc-tural member.

In addition, O temper is easier to bend for hull plating; and the result-ing temper—after factoring in strain hardening from rolling—leaves suffi-cient overload capacity in the finished hull components. With O temper, it’s likely that simple rolling of hull plat-ing can increase temper to, say, H116. But an already strain-hardened alloy, say H321, is significantly more diffi-cult to roll, and once rolled is strain-hardened possibly to a temper of H2 or more, depending on the curva-ture’s complexity. This leaves little capacity for deformation in the event of overloading.

Some old-school fabricators will apply the heat method common in steel fabrication to straighten a buck-led or bent aluminum plate. This is poor practice. Heat can significantly affect aluminum’s proof strength.

Even if the metal has been perfectly handled, cut, shaped, and fixed into position, it must still be welded. This process alters aluminum’s fatigue strength, sometimes dramatically if

Simple processes such as rolling and bending will work-harden the alloy, which in turn reduces its capacity for deformation and load-carrying. For

Shaping should be performed in the O temper (annealed) or H111 (another temper designation) alloys to avoid excessive localized strain-hardening.

Figure 5-2. Amount of Available “Overload” to Base Metal Proof Stress Limit of 5083 Alloy

% o

f Lo

ad O

ver

the

Bas

e P

roof

Str

ess 120

100

80

60

40

20

00 H2 H4 H6 H8 H9

Temper

Figure 5-1. Typical Properties of 5083 Alloy at Different Tempers

Stre

ss (

MPa

)

~330

240

140

Strain

Ultimate Tensile Strength

H2 temper

O temper

Figure 5. Typical Properties of 5083 Alloy at Different Tempers

Figure 6. Amount of Available “Overload” to Base Metal Proof Stress Limit of 5083 Alloy

and fabrication. And the cost of repairing a finished and flawed ves-sel can equal a large percentage of the initial build cost.

In steel construction, applying the correct filler wire or painting the bare metal surface to prevent obvious cor-rosion have become standard best practices, so ingrained and accepted than no one questions either the

need or the expense. This is QA at its best. By learning and adhering to similar rules and standard prac-tices, and recognizing that steel and aluminum require different approaches that should not be universally applied to both, the same high quality can be achieved in alumi-num construction.

—John Kecsmar

60 Professional BoatBuilder

correct procedures are not followed. Welding reduces fatigue properties even further than those shown in Figs. 2 and 4. When aluminum is immersed in a corrosive environment like seawater, its fatigue life is reduced yet again, as shown in Fig. 7. A welded joint does not have to be immersed in seawater for its fatigue life to be reduced. If it is wetted just once with seawater in a bilge or unventilated compartment, and then dries, the effect is the same: the salt deposits are deliquescent, meaning that in damp conditions the salt crys-tals attract water in the air to form a highly saline drop of liquid, probably more corrosive than seawater.

Fig. 7 illustrates that when alumi-num is welded, the reduction in fatigue strength “in air” is approxi-mately 50%, and when that weld is immersed in seawater, reduction can be as much as 90% of the static unwelded strength. So, static design limits appear to be rather meaningless for a welded aluminum boat that is constantly exposed to or immersed in seawater. To avert failures, designers must carefully position welded joints, and builders must practice compre-hensive quality control.

Frame Design Proper structural analysis is always

necessary in planning a complex structure. Incorrectly identifying load paths in connecting structure can cause serious problems in the framing for aluminum boats. When a typical stringer/longitudinal passes through a main transverse frame, cutouts are required so both structural members can be continuous. The size, shape, and connection detail of these cutouts requires care and an understanding of the true forces the frames will be sub-jected to. In Fig. 8, a typical arrange-ment shows an angle bar stringer, which could just as easily be a T or a bulb bar.

In the connection between the web of the transverse frame and the web of the longitudinal, seen in sec-tion, there is a 0.6" (15mm) radius. This prevents the weld of the trans-verse frame web (left to right on the page) from coming in contact with the weld of the web of the stringer. That’s important for two reasons: first, to avoid locating a weld over a weld (more on this later); second, the weld in the frame web carries a

Figure 6. S-N Curves of 5000 Series Weldments

Str

ess

(MPa)

160

140

120

100

80

60

40

20

0

Endurance (cycles)

100E+04 100E+05 100E+06 100E+07 100E+08 100E+09

In air

In seawater

Figure 7. S-N Curves of 5000-Series Weldments

Figure 7. Cutouts in Main Frame Web

63

Transverse frameR10(0.4")

R10(0.4")

R15(0.6")

Hull plate

Stringer

Figure 8. Cutouts in Main Frame Web

Figure 9. Poor Weld Return

June/July 2012 61

made crack waiting to propagate. Another common site for cracks is

the intersection of the flange of a lon-gitudinal and the web of a transverse frame. You can see in Fig. 8 that the flange of the angle bar has a defined gap between it and the frame web. Some designers and yards leave no gap, thinking that for a really stiff joint it’s often best to weld them together. This is incorrect. The result-ing joint is very stiff; however, the problem that manifests in service is similar to, but worse than, the biaxial stress riser on the lower-radius web connection, discussed above. The weld of the frame web is carrying shear transversely, and the flange in the longitudinal direction carries ten-sile load. The shear strength of alumi-num is much less than its tensile strength by a factor of √3; therefore, the allowable stress that the weld can take is reduced by 58% (1/√3)—the difference between the tensile and shear load capacity. Since the stringer

designer must ensure that the shear and bending loads are not excessive in any loading conditions in a sniped-end structural member.

Weld FlawsCracks in welds are the most com-

mon fatigue failure. They are generally the result of poor welding practice and not performing the standard checks to maintain quality. One of the most common is a simple start-stop crack initiated when too much heat is con-centrated in one location for too long, or not long enough to fully penetrate. Considered to be too time consuming, good welding practices are too often abandoned when a builder is under the pressure of production. A classic example is when a welder makes a long run and neglects to grind back the stop to prevent a void (Fig. 11). If the weld has gone cool or cold, a new weld run on top, or if performed too quickly, will rarely fuse deep into the root and so leaves a void—a ready-

load transversely, and the web of the stringer carries a load longitudinally. When these two welds join, the weld at the intersection is being “pulled” in two directions. If a load in each weld has a unit value of 1, then owing to the connection, the result-ing maximum load is the square root of the triangle of forces: √2, or 1.41. Thus, the weld at the intersection has an increase in stress of 41% over “as designed” stress. Overlay an increase of 41% in expected stress onto Fig. 7, and the resulting weakness and likely failure is apparent, commonly called a biaxial stress riser. (A triaxial is a joint in three directions). The size of this radius needs to ensure that the two welds don’t touch.

The weld on the frame web must also return around and be continu-ous. If the cutout is an odd shape/size, this too can increase the local stress as a discontinuity.

An example of a poor weld return is shown in Fig. 9. Weld returns are important because as the shear load in the weld runs out, the load has nothing to pass into. (FEA can estab-lish the structural load path to illus-trate the phenomenon.) But more importantly, the end of a weld always has too much heat without a runoff plate and creates a heat sink as the weld solidifies, and eventually leaves a classic “crater,” or start-stop, crack. This is visible in the poor joint shown in Fig. 9, as indicated by the circle. The arrow is pointing at the heat sink and small pore, a micro-scopic f law, which eventual ly becomes a crack that propagates under repeated loading. This type of flaw is also called a hot crack.

Shown in Fig. 10, a good return at the ends of a joint will prevent the heat sink, thus eliminating the flaw, and will provide a direct path around the joint to distribute the shear load when the vessel is in service.

If stringers are highly loaded, sniped ends will also exacerbate the fabrication problem. A snipe is the common method of finishing a struc-tural member if an end doesn’t attach to another member. For example, when an angle bar stops short of a frame and it’s not possible to attach a bracket to the frame, the bar is cut back at an angle of 30°. This gradu-ally reduces the change of section, thus lessening any stress concentra-tion from loads in the member. The

Figure 10. Good Weld Return

Figure 11. Undressed Weld Stops

62 Professional BoatBuilder

is constantly subjected to wave loads, it flexes with each passing wave in this extreme fatigue environment. As a result, the weld will crack, as shown in Fig. 12. Again, if the designer over-laid the calculated stress, with the √3 increase at this joint, onto the graph in Fig. 7, the implications for longev-ity are obvious. (While Fig. 7 shows a drop in strength at fatigued welds of roughly 90% in 5000-series aluminum, the results would be slightly worse in 6000 series.)

Poor fit-up is another common problem. If the plate is incorrectly sized, a large gap will result when it’s fitted up to the hull. Far too often, the welder will gap-fill to make it fit. A classic example of this is shown in Fig. 13, a cross section of the connec-tion of a stern tube passing through a waterjet duct, exhibiting a large void. The stern tube must be fitted to exact tolerances to ensure the correct gap for welding. Though the drawing called for just three weld beads, or passes—a root and two caps—numerous beads are visible on the finished and failed structure. This excessive gap-filling led to the inevitable lack of penetration (LOP) and lack of fusion (LOF). Inadequate penetration means the weld pool does not reach the weld root, and therefore a root gap remains. Finally, where the minimum cross-sectional area and heat-affected zone (HAZ) coincide, there’s the resulting fatigue crack at the weld toe (the line where the base metal meets the weld metal on the surface). The stern tube shown in Fig. 13 failed within a few months of going into service. Despite being a small joint, it was a very costly repair, because the vessel had to be put in dry dock. Poor training or just a fabricator’s momentary lapse in attention necessi-tated this expensive correction.

Apar t f rom being extremely unsightly, LOP/LOF seriously impacts the fatigue life of a welded joint, as illustrated in Fig. 14. That graph high-lights several other welding defects as well. As the void size increases, the number of cycles to failure decreases significantly, as does the correspond-ing fatigue strength with or without reinforcements. These unfused inter-faces between filler metal and base metal or between different layers of the filler material—characterized as fusion defects, or LOP and LOF—are difficult to detect with non-destructive

Figure 12. Improper Joint Weld

Figure 12. Lack of Penetration/Lack of Fusion

300

200

100

50

103 104 105 106 107

RR

RIFati

gue

Str

engt

h (M

Pa)

Cycles to Failure

Soundwelds

0.06"(1.5mm)

0.15"(3.8mm)

0.2"(5.3mm)

0.02"(0.5mm)

Size ofimperfectionand life range

RR = reinforcement removed RI = reinforcement intact

Figure 14. Lack of Penetration/Lack of Fusion

Figure 11. Excessive Gap Filling

Weld toe

Weld

Crack

LOP/LOF void

Figure 13. Excessive Gap Filling

D. K

OST

EAS

CO

URTESY

LLo

yd

’s R

eg

iste

R

June/July 2012 63

testing such as dye penetrant.Poor fit-up affects alignment as well

as gap filling. Misalignment can cause local ized stress concentrat ions because the resulting weld is over-sized, and also cause localized sec-ondary bending stresses because of discontinuity between butted plates. The combination of these localized effects exposes the weld toe to higher-than-expected stress (see Fig. 15).

The increase in secondary stresses is related to the geometric relationship of plate thickness and distance of mis-alignment. We can calculate the increase of stress at the weld toes in the plate by the simple relationship:

sN + sM = sN . (1 + 3.e/t)

Where:sN is axial stresst is plate thicknesse is eccentricity

The term in parentheses is the stress-magnification factor, Km. So if, for example, you have an axial mis-alignment of just 10% of the plate thickness (e/t = 0.1), Km = 1.3. A sim-ilar increase in stress results from a 1° angular misalignment, which means that if the plate is off by 10% and the angle of the joint is also rotated by just 1°, these slight imperfections combine for Km = 1 + 0.3 + 0.3 = 1.6. That’s a 60% increase of stress at the weld toe.

Fig. 16 shows the result of attempt-ing to overcome the gap where a frame rider butts into another rider running transversely. To fill the exces-sive gap, the welder “buttered” each edge with two passes, which cooled and locked in stress. The weld was finished by filling the middle with additional weld beads achieving no real penetration. Aluminum has approximately five times the thermal conductivity, and twice the rate of thermal expansion, of steel, so the heat from welding travels farther and faster than in steel. The crack is shown by the circles on either side of the rider, and the arrows indicate the hairline crack that has surfaced across the weld. It cracked because the over-sized volume of weld could not cool evenly across the whole joint, which caused internal thermal stress gradi-ents. Thus, it cracked as it cooled.

Fig. 17 illustrates several bracket -

Figure 13. Linear Misalignment

Str

ess

(MPa)

330

310

290

270

250100 20 30 40 50 60 70 80 90 100

Relative Eccentricity e/t (%)

AIMgSi1/S-AISi5t = 0.10" to 0.13" (2.5mm to 3.2mm)

e t

Figure 13. Linear Misalignment

Str

ess

(MPa)

330

310

290

270

250100 20 30 40 50 60 70 80 90 100

Relative Eccentricity e/t (%)

AIMgSi1/S-AISi5t = 0.10" to 0.13" (2.5mm to 3.2mm)

e t

Figure 15. Linear Misalignment

Figure 16. Crack in Oversized Weld

Figure 17. Bracket-Installation Errors

D. K

OST

EA

S A

ND

S.J. M

AD

DO

x

64 Professional BoatBuilder

installation errors. The crack circled in red is the most obvious one. It was caused by the lack of a proper weld return, which led to too much heat at the ends as the welder stopped, which in turn pulled the joint as it cooled. To the left of the crack, the welder gap-filled to overcome bad fit-up. You also can just see a hard saw cut, indicated by the arrow, in the bracket material. This hard edge will also crack.

Gaps and misalignment aside, over-sized or misshapen welds on their own are sources of local discontinu-ity. The stress in a typical weld varies from the nominal stress—what you expect the weld to carry—to a much higher peak at the weld toe. How this affects the fatigue life is shown in Fig. 18. Within a window of likely weld profiles, the lower the toe angle (meaning the more rollover), the lower the load the weld can carry. As the allowable stress is lower, so too is the fatigue life.

If we look at the profile of a butt weld as a discontinuity, it shortens the fatigue life of a structure. Fabricators can minimize the effect by dressing the weld bead, which reshapes the weld toe to eliminate rollover. The curves shown in Fig. 19 clearly illustrate that an as-welded joint, which shows the weld bead profile untouched, has a much lower load-carrying capacity, and thus a shorter fatigue life, than a dressed weld or the virgin plate.

Welding over a weld, and rewelding existing welds owing to mistakes and/or replacing plate, are far-too-common poor practices. When subjected to a typical UTS bend test, these joints show undesirable effects on the grain size. Fig. 20 shows an increase in grain size of 33% by the fourth repair. Hardness is proportional to yield strength, so as grain size increases, hardness decreases, which in turn reduces strength.

The effect of multiple welding repairs on mechanical strength is quantified in Fig. 21. As weld repairs increase, UTS consistently falls below accepted standards.

Weld temperatures in the heat-affected zone typically range from 932°F to 1,112°F (500°C to 600°C). At these temperatures, fine particles ca l led in termeta l l i cs—such as Al6(MnFe) or Al6Mn and Al3Fe— precipitate from the alloy, and are no longer soluble once the weld cools.

Figure 18. Effect of Repairs on Grain Size

Gra

in S

ize

(μm

)

95

90

85

80

75

70

65

60

55

500

(as welded)1 2 3 4 5

Number of Repairs

Root pass

Face pass

Figure 17. Effects of Weld Dressing on Fatigue

Fati

gue

Str

engt

h (M

Pa)

250

200

150

100

50

0103 4 6 4 6 4 6 2 25 5104 105 106 107 108

Cycles to Failure

Butt weld5000-series alloy

base metal

Rollover (over�ll)dressed �ush

as welded

Figure 16. Rollover (Over�ll) Toe Angle

Str

ess

Ran

ge (

MPa)

120

100

80

60

40

20

0120°100° 140° 160° 180°

Toe Angle α

Alloy NP5/6R = 0t = 0.37" (9.4mm)

toeα

Figure 18. Rollover (Overfill) Toe Angle

Figure 19. Effects of Weld Dressing on Fatigue

Figure 20. Effect of Repairs on Grain Size

ALu

min

um

An

d t

he s

eA

/PECH

INEY

RH

EN

ALU

HAN

DBO

OK

S. K

ATSA

S, J

. N

IKO

LAO

U, AN

D G

. PA

PA

DIM

ITRIO

U

June/July 2012 65

The increase in grain growth shown in Fig. 20 is consistent with the increase in intermetallics and porosity—sites of crack initiation—caused by repeated exposure to welding temper-atures. Welding over a weld increases the grain size and the number of intermetallics inside the alloy with each thermal cycle. This has the effect of greatly reducing strength and the fatigue life of the welded joint, and expanding the weakened area around a weld. The specific influence on fatigue life owing to the increase of intermetallics can be as much as a 30% reduction, and porosity alone can reduce the fatigue life by as much as 200 times, with increasing flaw size (Fig. 22). In repairs, the welder should always cut out the heat-affected zone to avoid the cumulative degradation of the metal from repeated welding.

Solutions In repairing the original failure

in our waterjet-powered vessel, we identified the fabrication faults that initiated fatigue-related failures in its complex stern structure. With FEA we found the locations of the load paths that the original design and fabrica-tion had failed to identify as sources of potential failures that could result in cracking (Fig. 23). By applying best practices in aluminum to the next generation of waterjet-powered ves-sels, we have avoided those faults in fabrication.

We introduced more transverse frames in the detail design, so we could then reduce or completely elim-inate longitudinals in the affected region. This provided easy access, so the welder could make one con-tinuous weld, thereby limiting the start-stops that introduce so much potential for cracking. With no longi-tudinals or need for cutouts through structural members, all the connections were smoothed, minimizing welding discontinuities.

During fabrication, builders now roll the plate with more care and attention, and along the grain rather than across the grain. Where feasi-ble, we introduced post-weld treat-ment to dress welds, reducing the stress concentrations at the weld toes (Fig. 24), which has been shown to improve the fatigue life.

The application of the new waterjet structure design and fabrication philoso-phies were fully adopted on a class of

Figure 21. Effects of Porosity

Cyc

les

to F

ract

ure

106

105

104

103

102

101

100

Porosity in % of Fracture Area

0 8 16 24 32

Respective increase of net section stressin the order of 50%

Factor of ~200on life

Butt weld with rolloverdressed �ush

Figure 19. Effects of Welding Repairs on Mechanical Strength

Ult

imat

e Te

nsile

Str

engt

h

290

280

270

260

250

240

230

220

210

200

ASM Handbook safety limit

Number of Repairs

1 2 430 (as welded)

Figure 21. Effects of Welding Repairs on Mechanical Strength

Figure 22. Effects of Porosity on Fatigue Life

Figure 23. Load Paths

66 Professional BoatBuilder

147.6' (45m), 200-metric-ton, 45-knot vessels. Twenty-three metric tons of water per second pass through the waterjet duct shown in Fig. 25. To the best of my knowledge, the boats have not experienced a single structural failure of the transom around the waterjets in their 17 years of service of 12–16 hour days, seven days a week. The same is true of the smaller repaired waterjet structure highlighted at the beginning of the article.

This evidence justifies the close attention to detail that took place during design and fabrication to ensure that a structure will be trouble-free for its lifetime. In all the waterjet-powered vessels I have designed during the past 17 years, I

Figure 25. Waterjet Duct

Figure 24. Post-weld Treatment of Flange-Welded Joint of Jet Duct

0.4" (10mm)machining allowance

Holes drilled inconjunction withKaMeWa �ange

Back face �ange

1"(25mm)

20.0°min

ø760

ø14

R8

Depth of grindingto be 0.02" (0.5mm) below bottom of anyvisible weld

Scale 1:1

0.05"(1.25mm)

0.04"(1mm)

Figure 24. Post-Weld Treatment of Flange-Welded Joint of Jet Duct

June/July 2012 67

have found that establishing correct procedures and sticking to them are the keys to a long-lived and warranty-free vessel.

About the Author: John Kecsmar formed the marine consultancy com-pany Ad hoc marine designs Ltd with nigel Warren in 2005, after spending nearly 20 years together at FBm Babcock marine, in newport, united Kingdom. John is on Lloyd’s Register technical Committee, RinA’s high speed technical committee, mCA’s high speed Advisory group, and snAme’s o-50 maritime Quality Culture group. he has designed high-speed aluminum vessels such as patrol boats, fast ferries, sWAths, and crew boats for more than 20 years and is the author of many technical papers on high-speed design, structural design, and fatigue. he lives in Japan.John dedicates this article to his friend and mentor, nigel Warren, who died during its preparation. John writes that nigel was very generous with his encyclopedic knowledge of boats, and is sadly missed.

For in-depth study of fatigue failures and aluminum boat construction, we include a list of the author’s technical papers and other sources he credits for this article.

Katsas, S., J. Nikolaou, G. Papadimitriou. “Microstructural changes accompany-ing repair welding in 5xxx aluminum alloys and their effect on the mechanical properties,” materials & design, Vol. 27, Issue 10, Elsevier, 2006.

Kecsmar, Fyfe, Hawkins, Shenoi, Price, Temarel, Read. “Fatigue Performance of Welded Aluminum Tee Connections,” FAST 1997.

Kecsmar, J. “Guidance on How Weld Quality Influences the Fatigue Life of Welded Aluminum Structures,” JASNAOE Conf., Vol. 3, Nov. 2006.

Kecsmar, J., R.A. Shenoi. “Some Notes on the Influence of Manufacturing on the Fatigue Life of Welded Aluminum Structures,” Journal of ship Production, Vol. 20, No. 3, SNAME, August 2004.

Kecsmar, J., N. Warren, J. Moore. “The evolution of advanced SLICE® Technology adapted to satisfy the HSC Code and commercial requirements,” RINA H.S.A.M.V. Conference, Shanghai, China, April 6, 2007.

Kecsmar, J., N. Warren, N. Sims. “Waterjet Propulsion: A Shipbuilder’s View,” RINA Conference, London, U.K., December 1–2, 1994.

Kosteas, D. “Fatigue Behaviour and Analysis,” Lecture 2401, Advanced Level, Training in Aluminum Application Technologies, European Aluminum Association, 1994.

Maddox, S.J. Fatigue strength of Welded structures. Abington Publishing, 2nd Ed, 1991, ISBN-10 1 85573 013 8.

Panico, J.R. “The effect of multiple weld repairs on the fatigue resistance of welded aluminum alloy 5083-O,” NTIS, Dec. 1979.

Pechiney Rhenalu Handbook. Aluminum and the sea, 1993. Polezhaeva, H., M. Malinowski. “Fatigue Strength of Aluminum Structural Details

of Special Service Craft,” FAST 2001, Southampton, U.K.

Further Reading