IPC-D-279 EN.pdf - 第48页

Solder , uniquely among the commonly used engineering metals, readily creeps and stress-relaxes at normal use tem- peratures; the rate of creep and stress-relaxation is highly temperature- and stress-level-dependent. Thu…

can be rather more complex [Ref. A-9: 9].

A-2.2 Global Expansion Mismatch

The global expansion mismatches result from differential

thermal expansions of an electronic component or connec-

tor and the printed board to which it is attached via the

surface mount solder joints. These thermal expansion dif-

ferences result from differences in the CTEs and thermal

gradients as the result of thermal energy being dissipated

within active components.

Global CTE-mismatches typically range from ∆α~2

ppm/°C (1 ppm=lx10

-6

) for CTE-tailored high reliability

assemblies to ~14 ppm/°C for ceramic components on

FR-4 printed boards. CTE-mismatches of ∆α <2 ppm/°C

are not achievable in reality as a consequence of the vari-

ability of the CTE values of the materials involved on both

components and printed boards.

Global thermal expansion mismatches typically are the

largest, since all three parameters determining the thermal

expansion mismatch—the CTE-mismatch, ∆α, the tem-

perature swing, ∆T, and the acting distance, L

—are large.

This global expansion mismatch will cyclically stress, and

thus fatigue, the solder joints. The cyclically cumulative

fatigue damage will ultimately cause the failure of one of

the solder joints, typically a corner joint, of the component

causing functional electrical failure that is initially intermit-

tent.

A-2.3 Local Expansion Mismatch

The local expansion mismatch results from differential

thermal expansions of the solder and the base material of

the electronic component or printed board to which it is

soldered. These thermal expansion differences result from

differences in the CTE of the solder and those of the base

materials together with thermal excursions [Refs. A-9: 4,7].

Local CTE-mismatches typically range from ∆α~7 ppm/°C

with copper to ~18 ppm/°C with ceramic and ~20 ppm/°C

with Alloy 42 and Kovar

. Local thermal expansion mis-

matches typically are smaller than the global expansion

mismatches, since the acting distance, the maximum wetted

area dimension, is much smaller—in the order of hundreds

of ~µm.

A-2.4 Internal Expansion Mismatch

An internal CTE-mismatch of ~6 ppm/°C results from the

different CTEs of the Sn-rich and Pb-rich phases of the

solder. Internal thermal expansion mismatches typically are

the smallest, since the acting distance, the size of the grain

structure, is much smaller than either the wetted length or

the component dimension—in the order of less than 25 µm

[Ref. A-9: 11].

A-2.5 Solder Attachment Failure

The failure of the solder attachment of a component to the

substrate to which it is surface mounted is commonly

deﬁned as the ﬁrst complete fracture of any of the solder

joints of which the component solder attachment consists.

Given that the loading of the solder joints is typically in

shear, rather than in tension, the mechanical failure of a

solder joint is not necessarily the same as the electrical

failure. Electrically, the mechanical failure of a solder joint

results, at least initially, in the occasional occurrence of a

short-duration (<1 µs) high-impedance event during either

a mechanical or thermal disturbance. From a practical point

of view, the solder joint failure is deﬁned as the ﬁrst obser-

vation of such an event.

For some applications this failure deﬁnition might be inad-

equate. For high-speed signals with sharp rise times signal

deterioration prior to the complete mechanical failure of a

solder joint might require a more stringent failure deﬁni-

tion. Similarly, for applications which subject the electronic

assemblies to signiﬁcant mechanical vibration and/or shock

loading, a failure deﬁnition that considers the mechanical

weakening of the solder joints as the result of the accumu-

lating fatigue damage might be necessary.

A-3.0 RELIABILITY PREDICTION MODELING

A-3.1 Creep-Fatigue Modeling

It has been experimentally shown [Refs. A-9: 2,4,12,13]

that the fatigue life of surface mount solder joints can be

described by a power law similar to the Coffin-Manson

low-cycle fatigue equation [Ref. A-9: 14] developed for

more typical engineering metals. For practical reasons and

as the direct consequence of the time-dependent stress-

relaxation/creep behavior of the solder at typical use envi-

ronments (see Table A-1), the specialized case of the

Coffin-Manson equation requires reversion to the more

general strain-energy relationship of Morrow [Ref. A-9:

15]; it also requires that the cyclic strain energy be based

on the total possible thermal expansion mismatch and that

the exponent is a function of temperature and time to pro-

vide a measure of the completeness of the stress-relaxation

process. The Engelmaier-Wild solder creep-fatigue equa-

tion [Refs. A-9: 1-6,9,12], subject to some caveats listed

later, relates the cyclic visco-plastic strain energy, repre-

sented by the cyclic fatigue damage term, ∆D, to the

median cyclic fatigue life for both isothermal-mechanical

and thermal cycling [Ref. A-9: 16]

(50%)=

[

2e’

∆D

]

−1

[Eq. A-1]

where

e’

= fatigue ductility coefficient, =0.325 for eutectic and

60/40 Sn/Pb solder (for other solders the value of e’

is expected to be somewhat different).

IPC-D-279 July 1996

Solder, uniquely among the commonly used engineering

metals, readily creeps and stress-relaxes at normal use tem-

peratures; the rate of creep and stress-relaxation is highly

temperature- and stress-level-dependent. Thus, the cyclic

fatigue damage term, ∆D, for practical reasons, has to be

based on the total potential damage at complete creep/

stress relaxation of the solder. For cyclic conditions that do

not allow sufficient time for complete stress relaxation ∆D

is larger than the actual fatigue damage. The temperature-

and time-dependent exponent, c, compensates for the

incomplete stress relaxation and is given by

c =−0.442 − 6x10

−4

+ 1.74x10

−2

ln(1 +

360

)

[Eq. A-2]

where

= mean cyclic solder joint temperature

= half-cycle dwell time in minutes.

The half-cycle dwell time relates to the cyclic frequency

and the shape of the cycles and represents the time avail-

able for the stress-relaxation/creep to take place.

In Eq. A-1 the exponent is given as (−1/c), which is men-

tally confusing; this format exists for historical reasons in

that the underlying work [Refs. A-9: 14,15] was always

stated this way. For typical electronic applications (T

to 100°C and t

= 15 to 720 minutes) the exponent (−1/c

ranges between 2.0 and 2.6.

Equations A-1 and A-2 come from a generic understanding

of the response of SM solder joints to cyclically accumu-

lating fatigue damage resulting from shear displacements

due to the global thermal expansion mismatches between

component and substrate. These shear displacements cause

time-independent yielding strains and time-, temperature-,

and stress-dependent creep/stress relaxation strains. These

strains, on a cyclic basis, form a visco-plastic strain energy

hysteresis loop which characterizes the solder joint

response to thermal cycling and whose area, given as the

damage term ∆D, is indicative of the cyclically accumulat-

ing fatigue damage. Hysteresis loops in the shear stress-

strain plane have been experimentally obtained [Refs. A-9:

13,17-19].

A-3.2 Damage Modeling

The assessment of the cyclically cumulating fatigue dam-

age is not a straight-forward task. While Eq. A-1 is widely

used, the question of how to best quantify the cyclic fatigue

damage is still hotly debated. The choices are primarily

between more complex ﬁnite-element analyses (FEA),

which can give more detailed information and can include

second-order effects, but require a large number of not

fully-supported assumptions [Ref. A-9: 20]; and closed-

form empirically-based relationships of the ﬁrst-order

design parameters, which cannot include second-order

effects and have use limitations due to their simple nature,

but allow, due to their simple form, a direct assessment of

the impact of the primary design parameters as well as

design trade-offs.

The following cyclic fatigue damage terms are of the sim-

pliﬁed closed-form type and should be utilized with the

application caveats that follow [Refs. A-9: 1-6,9,12,16,21].

The cyclic fatigue damage term for leadless SM solder

attachments, for which the stresses in the solder joints

exceed the solder yield strength and cause plastic yielding

of the solder, is

∆D(leadless)=

[

∆(α∆T)

]

[Eq. A-3]

For solder attachments with leads compliant enough, so

that the solder joint stresses are below the yield strength

and thus are not bounded by it, the cyclic fatigue damage

term is

∆D(leaded)=

[

∆(α ∆T)]

(919 kPa)Ah

]

[Eq. A-4]

where for English units the scaling coefficient is 133 psi

instead of 919 kPa.

Equation A-4 contains the design parameters that have a

ﬁrst-order inﬂuence on the reliability of SM solder attach-

ments. They are

A = effective minimum load bearing solder joint

area;

F = empirical ‘‘non-ideal’’ factor indicative of

deviations of real solder joints from idealizing

assumptions and accounting for secondary and

frequently intractable effects such as cyclic

warpage, cyclic transients, non-ideal solder

joint geometry, different solder crack propaga-

tion distances, brittle IMCs, Pb-rich boundary

layers, and solder/bonded-material expansion

differences, as well as inaccuracies and uncer-

tainties in the parameters in Eqs. A-1 through

A-4; 1.5>F> l.0 for ball/column-like leadless

solder joints (C4, C5, BGAs, CGAs),

1.2>F>0.7 for leadless solder joints with fillets

(castellated chip carriers and chip compo-

nents), F≈1 for solder attachments utilizing

compliant leads;

h = solder joint height;

= ‘‘diagonal’’ ﬂexural stiffness of unconstrained,

not soldered, corner-most component lead,

determined by strain methods [see Refs. A-9:

22-25] or FEA;

= maximum distance between component solder

joints measured from component solder joint

July 1996 IPC-D-279

pad centers, L

is sometimes referred to as the

distance from the neutral point (DNP);

= steady-state operating temperature for compo-

nent, substrate (T

for power dissipation in

component) during high temperature dwell;

C,0

S,O

= steady-state operating temperature for compo-

nent, substrate during low temperature dwell,

for non-operational (power off) half-cycles

C,O

S,0

;

= (1/4)(T

C,O

S,O

), mean cyclic solder

joint temperature;

= CTEs for component, substrate;

∆D = potential cyclic fatigue damage at complete

stress relaxation;

∆T

−T

C,O

, cyclic temperature swing for compo-

nent;

∆T

−T

S,O

, cycling temperature swing for sub-

strate (at component);

∆(α∆T) = ?α

∆T

−α

∆T

?, absolute cyclic expansion

mismatch, accounting for the effects of power

dissipation within the component as well as

temperature variations external to the compo-

nent;

∆α = ?α

−α

?, absolute difference in CTEs of com-

ponent and substrate, CTE-mismatch, because

of the inherent variability in material proper-

ties ∆α<2x10

-6

should not be used in calculat-

ing reliability.

A-3.3 CAVEAT 1 — Solder Joint Quality

The solder joint fatigue behavior and the resulting reliabil-

ity prediction equations. Eqs. A-1 through A-2, were deter-

mined from thermal cycling results of solder joints that

failed as a result of fracture of the solder, albeit sometimes

close to the IMC layers. For solder joints for which layered

structures are interposed between the base material and the

solder joints, these equations could be optimistic upper

bounds if the interposed layered structures become the

‘weakest link’ in the surface mount solder attachments.

Such layered structures could be: metallization layers that

have weak bonds to the underlying base material, or are

weak themselves, or dissolve essentially completely in the

solder; oxide or contamination layers preventing a proper

metallurgical bond of the solder to the underlying metal;

brittle IMC layers too thick due to too many or improperly

long high temperature processing steps.

Some material choices can lead to lower quality and

weaker solder joints because the material is more difficult

to wet and solder. The nickel/iron alloys, Kovar

and

Alloy 42, fall into this material category. The resulting

lower solder joint quality indicated in Table A-2 is also

clearly evident in Figure A-2, where the solder joint pull

strength is shown for a variety of differently prepared

Alloy 42 and copper leads. Alloy 42 leads, even when

etched or pre-ﬂowed at temperatures higher than can be

tolerated by the component, show a substantial reduction in

the solder joint pull strength relative to copper. In the worst

instance, the leads from one Alloy 42 manufacturer have a

pull strength of less than half of those with more typical

Alloy 42 and are essentially non-wettable.

Early failures of the solder attachments of components with

Alloy 42 lead frames and leads during accelerated testing

[Refs. A-9: 26-30] and manufacture [Ref. A-9: 31] have

been documented.

Solder joints which have solder joint heights (gaps) of

h<50 to 75µm also require special attention. For solder

joints that thin, the gap is essentially ﬁlled with intermetal-

lic compounds and those solder metals that do not go into

solution with the base metals to form the IMCs. Therefore

Eqs. A-1 and A-2 do not apply because these gaps are not

ﬁlled with solder [Ref. A-9: 32]. These materials do not

creep as readily, if at all, at the prevailing temperatures and

are typically more brittle, but much stronger than solder.

Thus, fatigue lives are longer than would be predicted from

Eqs. A-1 and A-2 unless overstress conditions occur.

On the other hand, the fatigue lives of solder attachments

can be underestimated by Eqs. A-1 through A-4 if the com-

ponent is underﬁlled with a load-bearing substance, e.g.,

epoxy [Ref. A-9: 33]. Components that are glued-down to

the substrate result in higher solder joint fatigue reliability,

since the solder joints are loaded in compression when the

adhesive contracts on cooling from the solder reﬂow tem-

peratures. Covercoats can either increase or decrease solder

joint fatigue lives depending on the properties of the cover-

coat and when and how it is applied. Parylene

has been

found to increase the solder joint fatigue life by about a

factor of three.

In general, caution might be indicated in all instances

where the predicted life is less than 1000 cycles, because

the severe loading conditions producing such short lives

are likely to produce different damage mechanisms or/and

failure modes.

A-3.4 CAVEAT 2 — Large Temperature Excursions

Solder joints experiencing large temperature swings

(−50°C to + 80°C) which extend both signiﬁcantly below

and signiﬁcantly above the temperature region bounded by

Table A−2 Quality of Solder Joints with Copper and

Alloy 42 Resulting from Different Reﬂow Temperatures

Reﬂow

Temperature

°C

Solder Joint Quality

60/40 Solder to

Copper

60/40 Solder to

Alloy 42

~210 just o.k. marginal to bad

~240 good just o.k.

~260 good good

IPC-D-279 July 1996